ASM Handbook: Volume 6: Welding, Brazing, and Soldering (Asm Handbook)

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VOLUME

ASM INTERNATIONAL

®

PUBLICATION INFORMATION AND CONTRIBUTORS

WELDING, BRAZING, AND SOLDERING WAS PUBLISHED IN 1993 AS VOLUME 6 OF THE ASM HANDBOOK. THE VOLUME WAS PREPARED UNDER THE DIRECTION OF THE ASM HANDBOOK COMMITTEE. VOLUME CHAIRMEN THE VOLUME CHAIRMEN WERE DAVID LEROY OLSON, THOMAS A. SIEWERT, STEPHEN LIU, AND GLEN R. EDWARDS. AUTHORS • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • • •

LAMET UFRGS BRUNO L. ALIA RICHARD L. ALLEY AMERICAN WELDING SOCIETY WILLIAM R. APBLETT, JR. WILLIAM A. BAESLACK III THE OHIO STATE UNIVERSITY WILLIAM BALLIS COLUMBIA GAS OF OHIO CLIFF C. BAMPTON ROCKWELL INTERNATIONAL SCIENCE CENTER PROBAL BANERJEE AUBURN UNIVERSITY JOHN G. BANKER EXPLOSIVE FABRICATORS INC. ROBERT G. BARTIFAY ALUMINUM COMPANY OF AMERICA ROY I. BATISTA ROY E. BEAL AMALGAMATED TECHNOLOGIES INC. RAYMOND E. BOHLMANN MCDONNELL AIRCRAFT COMPANY SÉRGIO D. BRANDI ESCOLA POLITECNICA DA USP JOHN A. BROOKS SANDIA NATIONAL LABORATORIES DONALD W. BUCHOLZ IBM FEDERAL SYSTEMS CORPORATION PAUL BURGARDT EG&G ROCKY FLATS PLANT ROGER A. BUSHEY THE ESAB GROUP INC. CHRIS CABLE FEIN POWER TOOL RICHARD D. CAMPBELL JOINING SERVICES INC. HOWARD CARY HOBART BROTHERS COMPANY HARVEY CASTNER EDISON WELDING INSTITUTE ALLEN CEDILOTE INDUSTRIAL TESTING LABORATORY SERVICES HARRY A. CHAMBERS TRW NELSON STUD WELDING C. CHRIS CHEN MICROALLOYING INTERNATIONAL INC. SHAOFENG CHEN AUBURN UNIVERSITY SHAO-PING CHEN LOS ALAMOS NATIONAL LABORATORY BRYAN A. CHIN AUBURN UNIVERSITY MICHAEL J. CIESLAK SANDIA NATIONAL LABORATORIES RODGER E. COOK THE WILKINSON COMPANY STEPHEN A. COUGHLIN ACF INDUSTRIES INC. MARK COWELL METCAL INC. RICHARD S. CREMISIO RESCORP INTERNATIONAL INC. CARL E. CROSS CRAIG DALLAM THE LINCOLN ELECTRIC COMPANY BRIAN DAMKROGER SANDIA NATIONAL LABORATORIES JOSEPH R. DAVIS DAVIS AND ASSOCIATES

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JANET DEVINE SONOBOND ULTRASONICS PAUL B. DICKERSON RAY DIXON LOS ALAMOS NATIONAL LABORATORY SUE DUNKERTON THE WELDING INSTITUTE KEVIN DUNN TEXAS INSTRUMENTS INC. CHUCK DVORAK UNI-HYDRO, INC. JIM DVORAK UNI-HYDRO, INC. ROBERT J. DYBAS GENERAL ELECTRIC COMPANY THOMAS W. EAGAR MASSACHUSETTS INSTITUTE OF TECHNOLOGY GLEN R. EDWARDS COLORADO SCHOOL OF MINES GRAHAM R. EDWARDS THE WELDING INSTITUTE W.H. ELLIOTT, JR. OAK RIDGE NATIONAL LABORATORY JOHN W. ELMER LAWRENCE LIVERMORE NATIONAL LABORATORY STEVEN C. ERNST EASTMAN CHEMICAL COMPANY WILLIAM FARRELL FERRANTI-SCIAKY COMPANY JOEL G. FELDSTEIN FOSTER WHEELER ENERGY CORPORATION DAVID A. FLEMING COLORADO SCHOOL OF MINES JAMES A. FORSTER TEXAS INSTRUMENTS INC. MICHAEL D. FREDERICKSON ELECTRONICS MANUFACTURING PRODUCTIVITY FACILITY EDWARD FRIEDMAN WESTINGHOUSE ELECTRIC CORPORATION R.H. FROST COLORADO SCHOOL OF MINES CHARLES E. FUERSTENAU LUCAS-MILHAUPT INC. EDWARD B. GEMPLER STANLEY S. GLICKSTEIN WESTINGHOUSE ELECTRIC CORPORATION JOHN A. GOLDAK CARLETON UNIVERSITY ROBIN GORDON EDISON WELDING INSTITUTE JERRY E. GOULD EDISON WELDING INSTITUTE JOHN B. GREAVES, JR. ELECTRONICS MANUFACTURING PRODUCTIVITY FACILITY F. JAMES GRIST JOHN F. GRUBB ALLEGHENY LUDLUM STEEL MAOSHI GU CARLETON UNIVERSITY IAN D. HARRIS EDISON WELDING INSTITUTE L.J. HART-SMITH DOUGLAS AIRCRAFT COMPANY DAN HAUSER EDISON WELDING INSTITUTE C.R. HEIPLE METALLURGICAL CONSULTANT HERBERT HERMAN STATE UNIVERSITY OF NEW YORK G. KEN HICKEN SANDIA NATIONAL LABORATORIES EVAN B. HINSHAW INCO ALLOYS INTERNATIONAL INC. D. BRUCE HOLLIDAY WESTINGHOUSE MARINE DIVISION S. IBARRA AMOCO CORPORATION J. ERNESTO INDACOCHEA UNIVERSITY OF ILLINOIS AT CHICAGO SUNIL JHA TEXAS INSTRUMENTS INC. JERALD E. JONES COLORADO SCHOOL OF MINES RAYMOND H. JUERS NAVAL SURFACE WARFARE CENTER WILLIAM R. KANNE, JR. WESTINGHOUSE SAVANNAH RIVER COMPANY MICHAEL J. KARAGOULIS GENERAL MOTORS CORPORATION MICHAEL KARAVOLIS TEXAS INSTRUMENTS INC. LENNART KARLSSON LULEÅ UNIVERSITY OF TECHNOLOGY MICHAEL E. KASSNER OREGON STATE UNIVERSITY DOUG D. KAUTZ LAWRENCE LIVERMORE NATIONAL LABORATORY W. DANIEL KAY WALL COLMONOY CORPORATION JAMES F. KEY IDAHO NATIONAL ENGINEERING LABORATORY H.-E. KIM SEOUL NATIONAL UNIVERSITY

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SAMUEL D. KISER INCO ALLOYS INTERNATIONAL INC. MARVIN L. KOHN FMC CORPORATION DAMIAN J. KOTECKI THE LINCOLN ELECTRIC COMPANY KENNETH KRYSIAC HERCULES INC. CHUCK LANDRY THERMAL DYNAMICS CHARLES LANE DURALCAN H.J. LATIMER TAYLOR-WINFIELD CORPORATION GLEN S. LAWRENCE FERRANTI-SCIAKY COMPANY KARL LAZAR WERNER LEHRHEUER FORSCHUNGSZENTRUM JÜLICH GMBH ALEXANDER LESNEWICH J.F. LIBSCH LEPEL CORPORATION TOM LIENERT THE OHIO STATE UNIVERSITY ALLEN C. LINGENFELTER LAWRENCE LIVERMORE NATIONAL LABORATORY DALE L. LINMAN CENTECH CORPORATION VONNE LINSE EDISON WELDING INSTITUTE JOHN C. LIPPOLD EDISON WELDING INSTITUTE JIAYAN LIU AUBURN UNIVERSITY STEPHEN LIU COLORADO SCHOOL OF MINES MATTHEW J. LUCAS, JR. GENERAL ELECTRIC COMPANY KEVIN A. LYTTLE PRAXAIR INC. KIM MAHIN SANDIA NATIONAL LABORATORIES MURRAY W. MAHONEY ROCKWELL INTERNATIONAL SCIENCE CENTER DARRELL MANENTE VAC-AERO INTERNATIONAL INC. RICHARD P. MARTUKANITZ PENNSYLVANIA STATE UNIVERSITY KOICHI MASUBUCHI MASSACHUSETTS INSTITUTE OF TECHNOLOGY DAVID K. MATLOCK COLORADO SCHOOL OF MINES R.B. MATTESON TAYLOR-WINFIELD CORPORATION STEVEN J. MATTHEWS HAYNES INTERNATIONAL INC. JYOTI MAZUMDER UNIVERSITY OF ILLINOIS AT URBANA-CHAMPAIGN C.N. MCCOWAN NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY KRIS MEEKINS LONG MANUFACTURING LTD. GREGORY MELEKIAN GENERAL MOTORS CORPORATION ANTHONY R. MELLINI, SR. MELLINI AND ASSOCIATES INC. DAVID W. MEYER THE ESAB GROUP INC. JULE MILLER HOWARD MIZUHARA WESGO INC. ARTHUR G. MOORHEAD OAK RIDGE NATIONAL LABORATORY MILO NANCE MARTIN MARIETTA ASTRONAUTICS GROUP E.D. NICHOLAS THE WELDING INSTITUTE DAVID NOBLE ARCO EXPLORATION AND PRODUCTION TECHNOLOGY THOMAS NORTH UNIVERSITY OF TORONTO DAVID B. O'DONNELL INCO ALLOYS INTERNATIONAL INC. JONATHAN S. OGBORN THE LINCOLN ELECTRIC COMPANY DAVID L. OLSON COLORADO SCHOOL OF MINES TOSHI OYAMA WESGO INC. R. ALAN PATTERSON LOS ALAMOS NATIONAL LABORATORY LARRY PERKINS WRIGHT LABORATORY DARYL PETER DARYL PETER AND ASSOCIATES MANFRED PETRI GERHARD PETRI GMBH & CO. KG DAVID H. PHILLIPS EDISON WELDING INSTITUTE ABE POLLACK MICROALLOYING INTERNATIONAL INC. BARRY POLLARD ANATOL RABINKIN ALLIEDSIGNAL AMORPHOUS METALS

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GEETHA RAMARATHNAM UNIVERSITY OF TORONTO EDWARD G. REINEKE EXPLOSIVE FABRICATORS INC. JULIAN ROBERTS THERMATOOL CORPORATION M. NED ROGERS BATESVILLE CASKET COMPANY J.R. ROPER EG&G ROCKY FLATS PLANT ROBERT S. ROSEN LAWRENCE LIVERMORE NATIONAL LABORATORY JAMES E. ROTH JAMES E. ROTH INC. WILLIAM J. RUPRECHT GENERAL ELECTRIC COMPANY K. SAMPATH CONCURRENT TECHNOLOGIES CORPORATION BERNARD E. SCHALTENBRAND ALUMINUM COMPANY OF AMERICA BERNARD SCHWARTZ NORFOLK SOUTHERN CORPORATION MEL M. SCHWARTZ SIKORSKY AIRCRAFT ANN SEVERIN LUCAS-MILHAUPT INC. THOMAS A. SIEWERT NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY HERSCHEL SMARTT IDAHO NATIONAL ENGINEERING LABORATORY RONALD B. SMITH ALLOY RODS CORPORATION WARREN F. SMITH THERMATOOL CORPORATION LANCE R. SOISSON WELDING CONSULTANTS INC. HARVEY D. SOLOMON GENERAL ELECTRIC COMPANY BRUCE R. SOMERS LEHIGH UNIVERSITY ROBERT E. SOMERS SOMERS CONSULTANTS ROGER K. STEELE AAR TECHNICAL CENTER FRANK STEIN TAYLOR-WINFIELD CORPORATION TIM STOTLER EDISON WELDING INSTITUTE ROBERT L. STROHL TWECO/ARCAIR ROBERT A. SULIT SULIT ENGINEERING VERN SUTTER AMERICAN WELDING INSTITUTE W.T. TACK MARTIN MARIETTA R. DAVID THOMAS, JR. R.D. THOMAS & COMPANY KARL THOMAS TECHNISCHE UNIVERSITÄT, BRAUNSCHWEIG RAYMOND G. THOMPSON UNIVERSITY OF ALABAMA AT BIRMINGHAM DONALD J. TILLACK D.J. TILLACK & ASSOCIATES CHON L. TSAI THE OHIO STATE UNIVERSITY SCHILLINGS TSANG EG&G ROCKY FLATS PLANT HENDRIKUS H. VANDERVELDT AMERICAN WELDING INSTITUTE RICCARDO VANZETTI VANZETTI SYSTEMS INC. PAUL T. VIANCO SANDIA NATIONAL LABORATORIES P. RAVI VISHNU LULEÅ UNIVERSITY OF TECHNOLOGY MARY B. VOLLARO UNIVERSITY OF CONNECTICUT A. WAHID COLORADO SCHOOL OF MINES DANIEL W. WALSH CALIFORNIA POLYTECHNIC STATE UNIVERSITY R. TERRENCE WEBSTER CONSULTANT JOHN R. WHALEN CONTOUR SAWS INC. NEVILLE T. WILLIAMS BRITISH STEEL FRED J. WINSOR WELDING CONSULTANT R. XU UNIVERSITY OF ILLINOIS AT CHICAGO XIAOSHU XU AMERICAN WELDING INSTITUTE PHILIP M. ZARROW SYNERGISTEK ASSOCIATES

REVIEWERS • • •

YONI ADONYI U.S. STEEL TECHNICAL CENTER RICHARD L. ALLEY AMERICAN WELDING SOCIETY BERNARD ALTSHULLER ALCAN INTERNATIONAL LTD.

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TED L. ANDERSON TEXAS A&M UNIVERSITY LLOYD ANDERSON MARION-INDRESCO INC. FRANK G. ARMAO ALCOA TECHNICAL CENTER DANIEL ARTHUR TELEDYNE MCKAY RICHARD E. AVERY NICKEL DEVELOPMENT INSTITUTE R.F. BACON TECUMSEH PRODUCTS COMPANY WALLY G. BADER WILLIAM A. BAESLACK III THE OHIO STATE UNIVERSITY CLIFF C. BAMPTON ROCKWELL INTERNATIONAL SCIENCE CENTER JOHN G. BANKER EXPLOSIVE FABRICATORS INC. GEORGE C. BARNES ROBERT G. BARTIFAY ALUMINUM COMPANY OF AMERICA ROY E. BEAL AMALGAMATED TECHNOLOGIES INC. GARY BECKA ALLIEDSIGNAL AEROSPACE COMPANY DAN BEESON EXXON PRODUCTION MALAYSIA DAVID M. BENETEAU CENTERLINE (WINDSOR) LTD. CHRISTOPHER C. BERNDT THE THERMAL SPRAY LABORATORY SURENDRA BHARGAVA GENERAL MOTORS INC. NORMAN C. BINKLEY EDISON WELDING INSTITUTE ROBERT A. BISHEL INCO ALLOYS INTERNATIONAL INC. R.A. BLACK BLACKS EQUIPMENT LTD. OMER W. BLODGETT THE LINCOLN ELECTRIC COMPANY RICHARD A. BRAINARD GENERAL DYNAMICS LAND SYSTEMS DIVISION GLENN H. BRAVE ASSOCIATION OF AMERICAN RAILROADS ROBERT S. BROWN CARPENTER TECHNOLOGY CORPORATION WILLIAM A. BRUCE EDISON WELDING INSTITUTE CHUCK CADDEN GENERAL MOTORS HARVEY R. CASTNER EDISON WELDING INSTITUTE ALLEN B. CEDILOTE INDUSTRIAL TESTING LABORATORY SERVICES CORPORATION KENNETH D. CHALLENGER SAN JOSE STATE UNIVERSITY P.R. CHIDAMBARAM COLORADO SCHOOL OF MINES BOB CHRISTOFFEL ROBIN CHURCHILL ESCO CORPORATION MICHAEL J. CIESLAK SANDIA NATIONAL LABORATORIES BRADLEY A. CLEVELAND MTS SYSTEMS CORPORATION NANCY C. COLE OAK RIDGE NATIONAL LABORATORY HAROLD R. CONAWAY ROCKWELL INTERNATIONAL RICHARD B. CORBIT GENERAL PUBLIC UTILITIES NUCLEAR CORPORATION MARK COWELL METCAL INC. NORM COX RESEARCH INC. JOHN A. CRAWFORD NAVAL SURFACE WARFARE CENTER DENNIS D. CROCKETT THE LINCOLN ELECTRIC COMPANY CARL E. CROSS NARENDRA B. DAHOTRE UNIVERSITY OF TENNESSEE SPACE INSTITUTE T. DEBROY PENNSYLVANIA STATE UNIVERSITY JOSEPH DEVITO THE ESAB GROUP INC. JOHN A. DEVORE GENERAL ELECTRIC COMPANY PAUL B. DICKERSON RAY DIXON LOS ALAMOS NATIONAL LABORATORY KARL E. DORSCHU WELDRING COMPANY INC. ROBERT J. DYBAS GENERAL ELECTRIC COMPANY THOMAS W. EAGAR MASSACHUSETTS INSTITUTE OF TECHNOLOGY BRUCE J. EBERHARD WESTINGHOUSE SAVANNAH RIVER COMPANY GLEN R. EDWARDS COLORADO SCHOOL OF MINES

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JOHN W. ELMER LAWRENCE LIVERMORE NATIONAL LABORATORY WERNER ENGELMAIER ENGELMAIER ASSOCIATES INC. CHRIS ENGLISH GE AIRCRAFT ENGINES ROBERT G. FAIRBANKS SCARROTT METALLURGICAL COMPANY HOWARD N. FARMER CONSULTANT DAVID A. FLEMING COLORADO SCHOOL OF MINES ROBERT FOLEY COLORADO SCHOOL OF MINES BOBBY FOLKENING FMC GROUND SYSTEMS DIVISION DARREL FREAR SANDIA NATIONAL LABORATORIES MICHAEL D. FREDERICKSON ELECTRONICS MANUFACTURING PRODUCTIVITY FACILITY EUGENE R. FREULER SOUDRONIC NEFTENBACH AG STEVEN A. GEDEON WELDING INSTITUTE OF CANADA BOB GIBBONS PLS MATERIALS INC. PAUL S. GILMAN ALLIEDSIGNAL INC. STANLEY S. GLICKSTEIN WESTINGHOUSE ELECTRIC CORPORATION JOHN A. GOLDAK CARLETON UNIVERSITY CARL GRAF EDISON WELDING INSTITUTE WILLIAM L. GREEN THE OHIO STATE UNIVERSITY CHUCK GREGOIRE NATIONAL STEEL CORPORATION ROBERT A. GRIMM EDISON WELDING INSTITUTE BRIAN GRINSELL THOMPSON WELDING INC. ROBIN GROSS-GOURLEY WESTINGHOUSE JOHN F. GRUBB ALLEGHENY LUDLUM STEEL BOB GUNOW, JR. VAC-MET INC. C. HOWARD HAMILTON WASHINGTON STATE UNIVERSITY JAMES R. HANNAHS PMI FOOD EQUIPMENT GROUP FRANK HANNEY ABCO WELDING & INDUSTRIAL SUPPLY INC. DAVID E. HARDT MASSACHUSETTS INSTITUTE OF TECHNOLOGY IAN D. HARRIS EDISON WELDING INSTITUTE MARK J. HATZENBELLER KRUEGER INTERNATIONAL DAN HAUSER EDISON WELDING INSTITUTE C.R. HEIPLE METALLURGICAL CONSULTANT J.S. HETHERINGTON HETHERINGTON INC. BARRY S. HEUER NOOTER CORPORATION ROGER B. HIRSCH UNITROL ELECTRONICS INC. TIM P. HIRTHE LUCAS-MILHAUPT HUGH B. HIX INTERNATIONAL EXPLOSIVE METALWORKING ASSOCIATION F. GALEN HODGE HAYNES INTERNATIONAL INC. RICHARD L. HOLDREN WELDING CONSULTANTS INC. ALAN B. HOPPER ROBERTSHAW TENNESSEE DIVISION CHARLES HUTCHINS C. HUTCHINS AND ASSOCIATES JENNIE S. HWANG IEM-FUSION INC. S. IBARRA AMOCO CORPORATION J. ERNESTO INDACOCHEA UNIVERSITY OF ILLINOIS AT CHICAGO GARY IRONS HOBART TAFA TECHNOLOGIES INC. JAMES R. JACHNA MODINE MANUFACTURING COMPANY ROBERT G. JAITE WOLFENDEN INDUSTRIES INC. JOHN C. JENKINS CONSULTANT KATHI JOHNSON HEXACON ELECTRIC COMPANY WILLIAM R. JONES VACUUM FURNACE SYSTEMS CORPORATION ROBERT W. JUD CHRYSLER CORPORATION WILLIAM F. KAUKLER UNIVERSITY OF ALABAMA IN HUNTSVILLE DOUG D. KAUTZ LAWRENCE LIVERMORE NATIONAL LABORATORY

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W. DANIEL KAY WALL COLMONOY CORPORATION JACQUE KENNEDY WESTINGHOUSE JAMES F. KING OAK RIDGE NATIONAL LABORATORY ANDREW G. KIRETA COPPER DEVELOPMENT ASSOCIATION INC. SAMUEL D. KISER INCO ALLOYS INTERNATIONAL INC. JOSEPH H. KISSEL ITT STANDARD FRED KOHLER CONSULTANT M.L. KOHN FMC CORPORATION DAMIAN J. KOTECKI THE LINCOLN ELECTRIC COMPANY SINDO KOU UNIVERSITY OF WISCONSIN-MADISON CURTIS W. KOVACH TECHNICAL MARKETING RESOURCES INC. LAWRENCE S. KRAMER MARTIN MARIETTA LABORATORIES RAYMOND B. KRIEGER AMERICAN CYANAMID COMPANY KENNETH KRYSIAC HERCULES INC. DANIEL KURUZAR MANUFACTURING TECHNOLOGY INC. RICHARD A. LAFAVE ELLIOTT COMPANY FRANK B. LAKE THE ESAB GROUP INC. JOHN D. LANDES UNIVERSITY OF TENNESSEE WERNER LEHRHEUER FORSCHUNGSZENTRUM JÜLICH GMBH J.F. LIBSCH LEPEL CORPORATION VONNE LINSE EDISON WELDING INSTITUTE JOHN C. LIPPOLD EDISON WELDING INSTITUTE STEPHEN LIU COLORADO SCHOOL OF MINES RONALD LOEHMAN ADVANCED MATERIALS LABORATORY PAUL T. LOVEJOY ALLEGHENY LUDLUM STEEL GEORGE LUCEY U.S. ARMY LABORATORY COMMAND KEVIN A. LYTTLE PRAXAIR INC. COLIN MACKAY MICROELECTRONICS AND COMPUTER TECHNOLOGY CORPORATION MICHAEL C. MAGUIRE SANDIA NATIONAL LABORATORIES KIM W. MAHIN SANDIA NATIONAL LABORATORIES WILLIAM E. MANCINI DUPONT DARRELL MANENTE VAC-AERO INTERNATIONAL INC. AUGUST F. MANZ A.F. MANZ ASSOCIATES RICHARD P. MARTUKANITZ PENNSYLVANIA STATE UNIVERSITY KOICHI MASUBUCHI MASSACHUSETTS INSTITUTE OF TECHNOLOGY STEVEN J. MATTHEWS HAYNES INTERNATIONAL JYOTI MAZUMDER UNIVERSITY OF ILLINOIS AT URBANA-CHAMPAIGN JIM MCMAHON DOALL COMPANY ALAN MEIER COLORADO SCHOOL OF MINES STANLEY MERRICK TELEDYNE MCKAY ROBERT W. MESSLER, JR. RENSSELAER POLYTECHNIC INSTITUTE E.A. METZBOWER U.S. NAVAL RESEARCH LABORATORY JOEL MILANO DAVID TAYLOR MODEL BASIN ROBERT A. MILLER SULZER PLASMA TECHNIK INC. HERBERT W. MISHLER EDISON WELDING INSTITUTE BRAJENDRA MISHRA COLORADO SCHOOL OF MINES HOWARD MIZUHARA WESGO INC. RICHARD MONTANA MID-FLORIDA TECHNICAL INSTITUTE JERRY MOODY WORLD WIDE WELDING RICHARD A. MORRIS NAVAL SURFACE WARFARE CENTER P.J. MUDGE THE WELDING INSTITUTE AMIYA MUKHERJEE UNIVERSITY OF CALIFORNIA BILL MYERS DRESSER-RAND INC.

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ERNEST F. NIPPES CONSULTANT DONG WON OH COLORADO SCHOOL OF MINES DAVID L. OLSON COLORADO SCHOOL OF MINES EDGAR D. OPPENHEIMER CONSULTING ENGINEER CARMEN PAPONETTI HI TECMETAL GROUP INC. MADHU PAREKH HOBART BROTHERS COMPANY SUBHASH R. PATI INTERNATIONAL PAPER COMPANY R. ALAN PATTERSON LOS ALAMOS NATIONAL LABORATORIES CHARLES C. PEASE CP METALLURGICAL ROBERT LEON PEASLEE WALL COLMONOY CORPORATION DARYL PETER DARYL PETER & ASSOCIATES LORENZ PFEIFER JOHN F. PFLZNIENSKI KOLENE CORPORATION DAVID H. PHILLIPS EDISON WELDING INSTITUTE EARL W. PICKERING, JR. CONSULTANT E.R. PIERRE CONSULTING WELDING ADVISOR JOHN PILLING MICHIGAN TECHNOLOGICAL UNIVERSITY ABE POLLACK MICROALLOYING INTERNATIONAL INC. BARRY POLLARD ALEXANDRE M. POPE COLORADO SCHOOL OF MINES JEFFREY W. POST J.W. POST & ASSOCIATES INC. TERRY PROFUGHI HI TECMETAL GROUP INC. ANATOL RABINKIN ALLIEDSIGNAL AMORPHOUS METALS JIM D. RABY SOLDERING TECH INTERNATIONAL TED RENSHAW CONSULTANT THERESA ROBERTS SIKAMA INTERNATIONAL DAVID E. ROBERTSON PACE INC. CHARLES ROBINO SANDIA NATIONAL LABORATORIES M.N. ROGERS ABB POWER T&D COMPANY INC. J.R. ROPER EG&G ROCKY FLATS PLANT N.V. ROSS AJAX MAGNETHERMIC DIETRICH K. ROTH ROMAN MANUFACTURING INC. JOHN RUFFING 3M FLUIDS LABORATORY WAYNE D. RUPERT ENGLEHARD CORPORATION J.D. RUSSELL THE WELDING INSTITUTE C.O. RUUD PENNSYLVANIA STATE UNIVERSITY EDMUND F. RYBICKI UNIVERSITY OF TULSA JONATHAN T. SALKIN ARC APPLICATIONS INC. MEL M. SCHWARTZ SIKORSKY AIRCRAFT JOE L. SCOTT DEVASCO INTERNATIONAL INC. ALAN P. SEIDLER RMI TITANIUM COMPANY OSCAR W. SETH CHICAGO BRIDGE & IRON COMPANY ANN SEVERIN LUCAS-MILHAUPT INC. LEWIS E. SHOEMAKER INCO ALLOYS INTERNATIONAL INC. LYNN E. SHOWALTER NEWPORT NEWS SHIPBUILDING THOMAS A. SIEWERT NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY ALLEN W. SINDEL BEGEMANN HEAVY INDUSTRIES INC. MICHAEL H. SKILLINGBERG REYNOLDS METALS COMPANY GERALD M. SLAUGHTER OAK RIDGE NATIONAL LABORATORY HERSCHEL SMARTT IDAHO NATIONAL ENGINEERING LABORATORY JAMES P. SNYDER II BETHLEHEM STEEL CORPORATION LANCE R. SOISSON WELDING CONSULTANTS INC. HARVEY D. SOLOMON GENERAL ELECTRIC BRUCE R. SOMERS LEHIGH UNIVERSITY

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NARASI SRIDHAR SOUTHWEST RESEARCH INSTITUTE BOB STANLEY NATIONAL TRAINING FUND ROGER K. STEELE ASSOCIATION OF AMERICAN RAILROADS ARCHIE STEVENSON MAGNESIUM ELEKRON INC. VIJAY K. STOKES GENERAL ELECTRIC TIM STOTLER EDISON WELDING INSTITUTE M.A. STREICHER CONSULTANT ROBERT L. STROHL TWECO/ARCAIR LAWRENCE STRYKER ALTECH INTERNATIONAL MARK TARBY WALL COLMONOY CORPORATION CLAY TAYLOR MERRICK AND COMPANY J.R. TERRILL CONSULTANT RAYMOND G. THOMPSON UNIVERSITY OF ALABAMA AT BIRMINGHAM J.S. THROWER GENERAL ELECTRIC POWER GENERATION DONALD J. TILLACK D.J. TILLACK & ASSOCIATES FELIX TOMEI TRUMPF INC. CHON L. TSAI THE OHIO STATE UNIVERSITY SCHILLINGS TSANG EG&G ROCKY FLATS PLANT M. NASIM UDDIN THYSSEN STEEL GROUP ELMAR UPITIS CBI TECHNICAL SERVICES COMPANY JAMES VAN DEN AVYLE SANDI NATIONAL LABORATORIES CLARENCE VAN DYKE LUCAS-MIHAUPT INC. HENDRIKUS H. VANDERVELDT AMERICAN WELDING INSTITUTE DAVID B. VEVERKA EDISON WELDING INSTITUTE PAUL T. VIANCO SANDIA NATIONAL LABORATORIES ROBERT G. VOLLMER R. WALLACH UNIVERSITY OF CAMBRIDGE SANDRA J. WALMSLEY WESTINGHOUSE ELECTRIC CORPORATION RICHARD A. WATSON THE P&LE CAR COMPANY CHRIS WEHLUS GENERAL MOTORS C.E.T. WHITE INDIUM CORPORATION OF AMERICA ROGER N. WILD ELLIOTT WILLNER LOCKHEED MISSILES & SPACE COMPANY RICHARD WILSON HOUSTON LIGHTING AND POWER COMPANY W.L. WINTERBOTTOM FORD MOTOR COMPANY A.P. WOODFIELD GENERAL ELECTRIC AIRCRAFT ENGINES JAMES B.C. WU STOODY COMPANY THOMAS ZACHARIA OAK RIDGE NATIONAL LABORATORY

FOREWORD COVERAGE OF JOINING TECHNOLOGIES IN THE ASM HANDBOOK HAS GROWN DRAMATICALLY OVER THE YEARS. A SHORT CHAPTER ON WELDING--EQUAL IN SIZE TO ABOUT 5 PAGES OF TODAY'S ASM HANDBOOK--APPEARED IN THE 1933 EDITION OF THE NATIONAL METALS HANDBOOK PUBLISHED BY THE AMERICAN SOCIETY OF STEEL TREATERS, ASM'S PREDECESSOR. THAT MATERIAL WAS EXPANDED TO 13 PAGES IN THE CLASSIC 1948 EDITION OF METALS HANDBOOK. THE FIRST FULL VOLUME ON WELDING AND BRAZING IN THE SERIES APPEARED IN 1971, WITH PUBLICATION OF VOLUME 6 OF THE 8TH EDITION OF METALS HANDBOOK. VOLUME 6 OF THE 9TH EDITION, PUBLISHED IN 1983, WAS EXPANDED TO INCLUDE COVERAGE OF SOLDERING. THE NEW VOLUME 6 OF THE ASM HANDBOOK BUILDS ON THE PROUD TRADITION ESTABLISHED BY THESE PREVIOUS VOLUMES, BUT IT ALSO REPRESENTS A BOLD NEW STEP FOR THE SERIES. THE HANDBOOK HAS NOT ONLY BEEN REVISED, BUT ALSO ENTIRELY

REFORMATTED TO MEET THE NEEDS OF TODAY'S MATERIALS COMMUNITY. OVER 90% OF THE ARTICLES IN THIS VOLUME ARE BRAND-NEW, AND THE REMAINDER HAVE BEEN SUBSTANTIALLY REVISED. MORE SPACE HAS BEEN DEVOTED TO COVERAGE OF SOLIDSTATE WELDING PROCESSES, MATERIALS SELECTION FOR JOINED ASSEMBLIES, WELDING IN SPECIAL ENVIRONMENTS, QUALITY CONTROL, AND MODELING OF JOINING PROCESSES, TO NAME BUT A FEW. INFORMATION ALSO HAS BEEN ADDED FOR THE FIRST TIME ABOUT JOINING OF SELECTED NONMETALLIC MATERIALS. WHILE A DELIBERATE ATTEMPT HAS BEEN MADE TO INCREASE THE AMOUNT OF CUTTINGEDGE INFORMATION PROVIDED, THE ORGANIZERS HAVE WORKED HARD TO ENSURE THAT THE HEART OF THE BOOK REMAINS PRACTICAL INFORMATION ABOUT JOINING PROCESSES, APPLICATIONS, AND MATERIALS WELDABILITY--THE TYPE OF INFORMATION THAT IS THE HALLMARK OF THE ASM HANDBOOK SERIES. PUTTING TOGETHER A VOLUME OF THIS MAGNITUDE IS AN ENORMOUS EFFORT AND COULD NOT HAVE BEEN ACCOMPLISHED WITHOUT THE DEDICATED AND TIRELESS EFFORTS OF THE VOLUME CHAIRPERSONS: DAVID L. OLSON, THOMAS A. SIEWERT, STEPHEN LIU, AND GLEN R. EDWARDS. SPECIAL THANKS ARE ALSO DUE TO THE SECTION CHAIRPERSONS, TO THE MEMBERS OF THE ASM HANDBOOK COMMITTEE, AND TO THE ASM EDITORIAL STAFF. WE ARE ESPECIALLY GRATEFUL TO THE OVER 400 AUTHORS AND REVIEWERS WHO HAVE CONTRIBUTED THEIR TIME AND EXPERTISE IN ORDER TO MAKE THIS HANDBOOK A TRULY OUTSTANDING INFORMATION RESOURCE. EDWARD H. KOTTCAMP, JR. PRESIDENT ASM INTERNATIONAL EDWARD L. LANGER MANAGING DIRECTOR ASM INTERNATIONAL PREFACE THE ASM HANDBOOK, VOLUME 6, WELDING, BRAZING, AND SOLDERING, HAS BEEN ORGANIZED INTO A UNIQUE FORMAT THAT WE BELIEVE WILL PROVIDE HANDBOOK USERS WITH READY ACCESS TO NEEDED MATERIALS-ORIENTED JOINING INFORMATION AT A MINIMAL LEVEL OF FRUSTRATION AND STUDY TIME. WHEN WE DEVELOPED THE ORGANIZATIONAL STRUCTURE FOR THIS VOLUME, WE RECOGNIZED THAT ENGINEERS, TECHNICIANS, RESEARCHERS, DESIGNERS, STUDENTS, AND TEACHERS DO NOT SEEK OUT JOINING INFORMATION WITH THE SAME LEVEL OF UNDERSTANDING, OR WITH THE SAME NEEDS. THEREFORE, WE ESTABLISHED DISTINCT SECTIONS THAT WERE INTENDED TO MEET THE SPECIFIC NEEDS OF PARTICULAR USERS. THE EXPERIENCED JOINING SPECIALIST CAN TURN TO THE SECTION "CONSUMABLE SELECTION, PROCEDURE DEVELOPMENT, AND PRACTICE CONSIDERATIONS" AND FIND DETAILED JOINING MATERIALS DATA ON A WELL-DEFINED PROBLEM. THIS HANDBOOK ALSO PROVIDES GUIDANCE FOR THOSE WHO NOT ONLY MUST SPECIFY THE JOINING PRACTICE, BUT ALSO THE MATERIALS TO BE JOINED. THE SECTION "MATERIALS SELECTION FOR JOINED ASSEMBLIES" CONTAINS COMPREHENSIVE INFORMATION ABOUT THE PROPERTIES, APPLICATIONS, AND WELDABILITIES OF THE MAJOR CLASSES OF STRUCTURAL MATERIALS. TOGETHER, THESE TWO MAJOR SECTIONS OF THE HANDBOOK SHOULD PROVIDE AN ENGINEER ASSIGNED A LOOSELY DEFINED DESIGN PROBLEM WITH THE MEANS TO MAKE INTELLIGENT CHOICES FOR COMPLETING AN ASSEMBLY. FREQUENTLY, TECHNOLOGISTS ARE CALLED UPON TO INITIATE AND ADOPT WELDING PROCESSES WITHOUT IN-DEPTH KNOWLEDGE OF THESE PROCESSES OR THE SCIENTIFIC

PRINCIPLES THAT IMPACT THE PROPERTIES AND PERFORMANCE OF WELDMENTS. THE SECTIONS "FUNDAMENTALS OF JOINING" AND "JOINING PROCESSES" ARE DESIGNED TO MEET THE NEEDS OF THESE USERS, OR ANYONE WHO NEEDS BASIC BACKGROUND INFORMATION ABOUT JOINING PROCESSES AND PRINCIPLES. WELDING, BRAZING, AND SOLDERING ARE TRULY INTERDISCIPLINARY ENTERPRISES; NO INDIVIDUAL CAN BE EXPECTED TO BE AN EXPERT IN ALL ASPECTS OF THESE TECHNOLOGIES. THEREFORE, WE HAVE ATTEMPTED TO PROVIDE A HANDBOOK THAT CAN BE USED AS A COMPREHENSIVE REFERENCE BY ANYONE NEEDING MATERIALS-RELATED JOINING INFORMATION. MANY COLLEAGUES AND FRIENDS CONTRIBUTED THEIR TIME AND EXPERTISE TO THIS HANDBOOK, AND WE ARE VERY GRATEFUL FOR THEIR EFFORTS. WE WOULD ALSO LIKE TO EXPRESS OUR THANKS TO THE AMERICAN WELDING SOCIETY FOR THEIR COOPERATION AND ASSISTANCE IN THIS ENDEAVOR. DAVID LEROY OLSON, COLORADO SCHOOL OF MINES THOMAS A. SIEWERT, NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY STEPHEN LIU, COLORADO SCHOOL OF MINES GLEN R. EDWARDS, COLORADO SCHOOL OF MINES OFFICERS AND TRUSTEES OF ASM INTERNATIONAL (1992-1993)

OFFICERS • • • • •

EDWARD H. KOTTCAMP, JR. PRESIDENT AND TRUSTEE SPS TECHNOLOGIES JACK G. SIMON VICE PRESIDENT AND TRUSTEE GENERAL MOTORS CORPORATION WILLIAM P. KOSTER IMMEDIATE PAST PRESIDENT AND TRUSTEE METCUT RESEARCH ASSOCIATES, INC. EDWARD L. LANGER SECRETARY AND MANAGING DIRECTOR ASM INTERNATIONAL LEO G. THOMPSON TREASURER LINDBERG CORPORATION

TRUSTEES • • • • • • • • •

WILLIAM H. ERICKSON FDP ENGINEERING NORMAN A. GJOSTEIN FORD MOTOR COMPANY NICHOLAS C. JESSEN, JR. MARTIN MARIETTA ENERGY SYSTEMS, INC. E. GEORGE KENDALL NORTHROP AIRCRAFT GEORGE KRAUSS COLORADO SCHOOL OF MINES LYLE H. SCHWARTZ NATIONAL INSTITUTE OF STANDARDS & TECHNOLOGY GERNANT E. MAURER SPECIAL METALS CORPORATION ALTON D. ROMIG, JR. SANDIA NATIONAL LABORATORIES MERLE L. THORPE HOBART TAFA TECHNOLOGIES, INC.

MEMBERS OF THE ASM HANDBOOK COMMITTEE (1992-1993) • • • •

ROGER J. AUSTIN (CHAIRMAN 1992-; MEMBER 1984-) CONCEPT SUPPORT AND DEVELOPMENT CORPORATION DAVID V. NEFF (VICE CHAIRMAN 1992-; MEMBER 1986-) METAULLICS SYSTEMS TED L. ANDERSON (1991-) TEXAS A&M UNIVERSITY BRUCE P. BARDES (1993-) MIAMI UNIVERSITY

• • • • • • • • • • • • • • • • • • • • • •

ROBERT J. BARNHURST (1988-) NORANDA TECHNOLOGY CENTRE TONI BRUGGER (1993-) PHOENIX PIPE & TUBE COMPANY STEPHEN J. BURDEN (1989-) CRAIG V. DARRAGH (1989-) THE TIMKEN COMPANY RUSSELL J. DIEFENDORF (1990-) CLEMSON UNIVERSITY AICHA EISHABINI-RIAD (1990-) VIRGINIA POLYTECHNIC & STATE UNIVERSITY GREGORY A. FETT (1993-) DANA CORPORATION MICHELLE M. GAUTHIER (1990-) RAYTHEON COMPANY TONI GROBSTEIN (1990-) NASA LEWIS RESEARCH CENTER SUSAN HOUSH (1990-) DOW CHEMICAL U.S.A. DENNIS D. HUFFMAN (1982-) THE TIMKEN COMPANY S. JIM LBARRA (1991-) AMOCO RESEARCH CENTER J. ERNESTO INDACOCHEA (1987-) UNIVERSITY OF ILLINOIS AT CHICAGO PETER W. LEE (1990-) THE TIMKEN COMPANY WILLIAM L. MANKINS (1989-) INCO ALLOYS INTERNATIONAL, INC. RICHARD E. ROBERTSON (1990-) UNIVERSITY OF MICHIGAN JOGENDER SINGH (1993-) NASA GEORGE C. MARSHALL SPACE FLIGHT CENTER JEREMY C. ST. PIERRE (1990-) HAYES HEAT TREATING CORPORATION EPHRAIM SUHIR (1990-) AT&T BELL LABORATORIES KENNETH TATOR (1991-) KTA-TATOR, INC. MALCOLM THOMAS (1993-) ALLISON GAS TURBINES WILLIAM B. YOUNG (1991-) DANA CORPORATION

PREVIOUS CHAIRMEN OF THE ASM HANDBOOK COMMITTEE • • • • • • • • • • • • • • • • • • • • • • • •

R.S. ARCHER (1940-1942) (MEMBER 1937-1942) L.B. CASE (1931-1933) (MEMBER 1927-1933) T.D. COOPER (1984-1986) (MEMBER 1981-1986) E.O. DIXON (1952-1954) (MEMBER 1947-1955) R.L. DOWDELL (1938-1939) (MEMBER 1935-1939) J.P. GILL (1937) (MEMBER 1934-1937) J.D. GRAHAM (1966-1968) (MEMBER 1961-1970) J.F. HARPER (1923-1926) (MEMBER 1923-1926) C.H. HERTY, JR. (1934-1936) (MEMBER 1930-1936) D.D. HUFFMAN (1986-1990) (MEMBER 1982-1990) J.B. JOHNSON (1948-1951) (MEMBER 1944-1951) L.J. KORB (1983) (MEMBER 1978-1983) R.W.E. LEITER (1962-1963) (MEMBER 1955-1958, 1960-1964) G.V. LUERSSEN (1943-1947) (MEMBER 1942-1947) G.N. MANIAR (1979-1980) (MEMBER 1974-1980) J.L. MCCALL (1982) (MEMBER 1977-1982) W.J. MERTEN (1927-1930) (MEMBER 1923-1933) D.L. OLSON (1990-1992) (MEMBER 1982-1988, 1989-1992) N.E. PROMISEL (1955-1961) (MEMBER 1954-1963) G.J. SHUBAT (1973-1975) (MEMBER 1966-1975) W.A. STADTLER (1969-1972) (MEMBER 1962-1972) R. WARD (1976-1978) (MEMBER 1972-1978) M.G.H. WELLS (1981) (MEMBER 1976-1981) D.J. WRIGHT (1964-1965) (MEMBER 1959-1967)

STAFF ASM INTERNATIONAL STAFF WHO CONTRIBUTED TO THE DEVELOPMENT OF THE VOLUME INCLUDED WILLIAM W. SCOTT, JR., DIRECTOR OF TECHNICAL PUBLICATIONS; SCOTT D.

HENRY, MANAGER OF HANDBOOK DEVELOPMENT; SUZANNE E. HAMPSON, PRODUCTION PROJECT MANAGER; THEODORE B. ZORC, TECHNICAL EDITOR; FAITH REIDENBACH, CHIEF COPY EDITOR; LAURIE A. HARRISON, EDITORIAL ASSISTANT; NANCY M. SOBIE, PRODUCTION ASSISTANT. EDITORIAL ASSISTANCE WAS PROVIDED BY JOSEPH R. DAVIS, KELLY FERJUTZ, NIKKI D. WHEATON, AND MARA S. WOODS. CONVERSION TO ELECTRONIC FILES ASM HANDBOOK, VOLUME 6, WELDING, BRAZING, AND SOLDERING WAS CONVERTED TO ELECTRONIC FILES IN 1998. THE CONVERSION WAS BASED ON THE SECOND PRINTING (1994). NO SUBSTANTIVE CHANGES WERE MADE TO THE CONTENT OF THE VOLUME, BUT SOME MINOR CORRECTIONS AND CLARIFICATIONS WERE MADE AS NEEDED. ASM INTERNATIONAL STAFF WHO CONTRIBUTED TO THE CONVERSION OF THE VOLUME INCLUDED SALLY FAHRENHOLZ-MANN, BONNIE SANDERS, SCOTT HENRY, ROBERT BRADDOCK, AND MARLENE SEUFFERT. THE ELECTRONIC VERSION WAS PREPARED UNDER THE DIRECTION OF WILLIAM W. SCOTT, JR., TECHNICAL DIRECTOR, AND MICHAEL J. DEHAEMER, MANAGING DIRECTOR. COPYRIGHT INFORMATION (FOR PRINT VOLUME) COPYRIGHT © 1993 BY ASM INTERNATIONAL ALL RIGHTS RESERVED. ASM HANDBOOK IS A COLLECTIVE EFFORT INVOLVING THOUSANDS OF TECHNICAL SPECIALISTS. IT BRINGS TOGETHER IN ONE BOOK A WEALTH OF INFORMATION FROM WORLD-WIDE SOURCES TO HELP SCIENTISTS, ENGINEERS, AND TECHNICIANS SOLVE CURRENT AND LONG-RANGE PROBLEMS. GREAT CARE IS TAKEN IN THE COMPILATION AND PRODUCTION OF THIS VOLUME, BUT IT SHOULD BE MADE CLEAR THAT NO WARRANTIES, EXPRESS OR IMPLIED, ARE GIVEN IN CONNECTION WITH THE ACCURACY OR COMPLETENESS OF THIS PUBLICATION, AND NO RESPONSIBILITY CAN BE TAKEN FOR ANY CLAIMS THAT MAY ARISE. NOTHING CONTAINED IN THE ASM HANDBOOK SHALL BE CONSTRUED AS A GRANT OF ANY RIGHT OF MANUFACTURE, SALE, USE, OR REPRODUCTION, IN CONNECTION WITH ANY METHOD, PROCESS, APPARATUS, PRODUCT, COMPOSITION, OR SYSTEM, WHETHER OR NOT COVERED BY LETTERS PATENT, COPYRIGHT, OR TRADEMARK, AND NOTHING CONTAINED IN THE ASM HANDBOOK SHALL BE CONSTRUED AS A DEFENSE AGAINST ANY ALLEGED INFRINGEMENT OF LETTERS PATENT, COPYRIGHT, OR TRADEMARK, OR AS A DEFENSE AGAINST LIABILITY FOR SUCH INFRINGEMENT. COMMENTS, CRITICISMS, AND SUGGESTIONS ARE INVITED, AND SHOULD BE FORWARDED TO ASM INTERNATIONAL. LIBRARY OF CONGRESS CATALOGING-IN-PUBLICATION DATA (FOR PRINT VOLUME) ASM HANDBOOK (REVISED VOL. 6) METALS HANDBOOK. VOLS. 1-2 HAVE TITLE: METALS HANDBOOK. VOL. 4 LACKS ED. STATEMENTS. INCLUDES BIBLIOGRAPHICAL REFERENCES AND INDEXES. CONTENTS: V. 1. PROPERTIES AND SELECTION-IRONS, STEELS, AND HIGH-PERFORMANCE ALLOYS-V. 2. PROPERTIES AND SELECTION-NONFERROUS ALLOYS

AND SPECIAL-PURPOSE MATERIALS-[ETC.]-V. 6. WELDING, BRAZING, AND SOLDERING. 1. METALS-HANDBOOKS, MANUALS, ETC. 2. METAL-WORK-HANDBOOKS, MANUALS, ETC. I. ASM INTERNATIONAL. HANDBOOK COMMITTEE. II. TITLE: METALS HANDBOOK. TA459.M43 1990 620.1'6 90-115 ISBN 0-87170-377-7(V.1) SAN 204-7586 ISBN 0-87170-382-3 PRINTED IN THE UNITED STATES OF AMERICA Energy Sources Used for Fusion Welding Thomas W. Eagar, Massachusetts Institute of Technology

Introduction WELDING AND JOINING processes are essential for the development of virtually every manufactured product. However, these processes often appear to consume greater fractions of the product cost and to create more of the production difficulties than might be expected. There are a number of reasons that explain this situation. First, welding and joining are multifaceted, both in terms of process variations (such as fastening, adhesive bonding, soldering, brazing, arc welding, diffusion bonding, and resistance welding) and in the disciplines needed for problem solving (such as mechanics, materials science, physics, chemistry, and electronics). An engineer with unusually broad and deep training is required to bring these disciplines together and to apply them effectively to a variety of processes. Second, welding or joining difficulties usually occur far into the manufacturing process, where the relative value of scrapped parts is high. Third, a very large percentage of product failures occur at joints because they are usually located at the highest stress points of an assembly and are therefore the weakest parts of that assembly. Careful attention to the joining processes can produce great rewards in manufacturing economy and product reliability. The Section "Fusion Welding Processes" in this Volume provides details about equipment and systems for the major fusion welding processes. The purpose of this Section of the Volume is to discuss the fundamentals of fusion welding processes, with an emphasis on the underlying scientific principles. Because there are many fusion welding processes, one of the greatest difficulties for the manufacturing engineer is to determine which process will produce acceptable properties at the lowest cost. There are no simple answers. Any change in the part geometry, material, value of the end product, or size of the production run, as well as the availability of joining equipment, can influence the choice of joining method. For small lots of complex parts, fastening may be preferable to welding, whereas for long production runs, welds can be stronger and less expensive. The perfect joint is indistinguishable from the material surrounding it. Although some processes, such as diffusion bonding, can achieve results that are very close to this ideal, they are either expensive or restricted to use with just a few materials. There is no universal process that performs adequately on all materials in all geometries. Nevertheless, virtually any material can be joined in some way, although joint properties equal to those of the bulk material cannot always be achieved. The economics of joining a material may limit its usefulness. For example, aluminum is used extensively in aircraft manufacturing and can be joined by using adhesives or fasteners, or by welding. However, none of these processes has proven economical enough to allow the extensive replacement of steel by aluminum in the frames of automobiles. An increased use of composites in aircrafts is limited by an inability to achieve adequate joint strength.

It is essential that the manufacturing engineer work with the designer from the point of product conception to ensure that compatible materials, processes, and properties are selected for the final assembly. Often, the designer leaves the problem of joining the parts to the manufacturing engineer. This can cause an escalation in cost and a decrease in reliability. If the design has been planned carefully and the parts have been produced accurately, the joining process becomes much easier and cheaper, and both the quality and reliability of the product are enhanced. Generally, any two solids will bond if their surfaces are brought into intimate contact. One factor that generally inhibits this contact is surface contamination. Any freshly produced surface exposed to the atmosphere will absorb oxygen, water vapor, carbon dioxide, and hydrocarbons very rapidly. If it is assumed that each molecule that hits the surface will be absorbed, then the time-pressure value to produce a monolayer of contamination is approximately 0.001 Pa · s (10-8 atm · s). For example, at a pressure of 1 Pa (10-5 atm), the contamination time is 10-3 s, whereas at 0.1 MPa (1 atm), it is only 10 × 10-9 s. In fusion welding, intimate interfacial contact is achieved by interposing a liquid of substantially similar composition as the base metal. If the surface contamination is soluble, then it is dissolved in the liquid. If it is insoluble, then it will float away from the liquid-solid interface. Energy Sources Used for Fusion Welding Thomas W. Eagar, Massachusetts Institute of Technology

Energy-Source Intensity One distinguishing feature of all fusion welding processes is the intensity of the heat source used to melt the liquid. Virtually every concentrated heat source has been applied to the welding process. However, many of the characteristics of each type of heat source are determined by its intensity. For example, when considering a planar heat source diffusing into a very thick slab, the surface temperature will be a function of both the surface power density and the time. Figure 1 shows how this temperature will vary on steel with power densities that range from 400 to 8000 W/cm2. At the lower value, it takes 2 min to melt the surface. If that heat source were a point on the flat surface, then the heat flow would be divergent and might not melt the steel. Rather, the solid metal would be able to conduct away the heat as fast as it was being introduced. It is generally found that heat-source power densities of approximately 1000 W/cm2 are necessary to melt most metals.

FIG. 1 TEMPERATURE DISTRIBUTION AFTER A SPECIFIC HEATING TIME IN A THICK STEEL PLATE HEATED

UNIFORMLY ON ONE SURFACE AS A FUNCTION OF APPLIED HEAT INTENSITY; INITIAL TEMPERATURE OF PLATE IS 25 °C (77 °F)

At the other end of the power-density spectrum, heat intensities of 106 or 107 W/cm2 will vaporize most metals within a few microsecond. At levels above these values, all of the solid that interacts with the heat source will be vaporized, and no fusion welding can occur. Thus, the heat sources for all fusion welding processes should have power densities between approximately 0.001 and 1 MW/cm2. This power-density spectrum is shown in Fig. 2, along with the points at which common joining processes are employed.

FIG. 2 SPECTRUM OF PRACTICAL HEAT INTENSITIES USED FOR FUSION WELDING

The fact that power density is inversely related to the interaction time of the heat source on the material is evident in Fig. 1. Because this represents a transient heat conduction problem, one can expect the heat to diffuse into the steel to a depth that increases as the square root of time, that is, from the Einstein equation: X ~ αt

(EQ 1)

where x is the distance that the heat diffuses into the solid, in centimeters: α is the thermal diffusivity of the solid, in cm2/s; and t is the time in seconds. Tables 1 and 2 give the thermal diffusivities of common elements and common alloys, respectively.

TABLE 1 THERMAL DIFFUSIVITIES OF COMMON ELEMENTS FROM 20 TO 100 °C (68 TO 212 °F)

ELEMENT

ALUMINUM ANTIMONY BERYLLIUM BISMUTH CADMIUM CARBON COBALT COPPER GALLIUM GERMANIUM GOLD HAFNIUM

DENSITY

HEAT CAPACITY

THERMAL CONDUCTIVITY

g/cm3

lb/in.3

j/kg · k

calit/g · °c

w/m · k

calit/cm · s · °c

2.699 6.62 1.848 9.80 8.65 2.25 8.85 8.96 5.907 5.323 19.32 13.09

0.098 0.239 0.067 0.354 0.313 0.081 0.320 0.324 0.213 0.192 0.698 0.472

900 205 1880 123 230 691 414 385 331 306 131 147

0.215 0.049 0.45 0.0294 0.055 0.165 0.099 0.092 0.079 0.073 0.0312 0.0351

221 19 147 8 92 24 69 394 29-38 59 297 22

0.53 0.045 0.35 0.020 0.22 0.057 0.165 0.941 0.07-0.09 0.14 0.71 0.053

mm2/s THERMAL DIFFUSIVITY cm2/s 91 0.91 14 0.14 42 0.42 7 0.09 46 0.46 15 0.15 19 0.188 114 1.14 17 0.17 36 0.36 118 1.178 12 0.12

INDIUM IRIDIUM IRON LEAD MAGNESIUM MOLYBDENUM NICKEL NIOBIUM PALLADIUM PLATINUM PLUTONIUM RHODIUM SILICON SILVER SODIUM TANTALUM TIN TITANIUM TUNGSTEN URANIUM VANADIUM ZINC ZIRCONIUM

7.31 22.5 7.87 11.36 1.74 10.22 8.902 8.57 12.02 21.45 19.84 12.44 2.33 10.49 0.9712 16.6 7.2984 4.507 19.3 19.07 6.1 7.133 6.489

0.264 0.813 0.284 0.410 0.063 0.369 0.322 0.310 0.434 0.775 0.717 0.449 0.084 0.379 0.035 0.600 0.264 0.163 0.697 0.689 0.22 0.258 0.234

239 129 460 129 1025 276 440 268 244 131 138 247 678 234 1235 142 226 519 138 117 498 383 280

0.057 0.0307 0.11 0.0309 0.245 0.066 0.105 0.064 0.0584 0.0314 0.033 0.059 0.162 0.0559 0.295 0.034 0.054 0.124 0.033 0.0279 0.119 0.0915 0.067

24 59 75 35 154 142 92 54 70 69 8 88 84 418 134 54 63 22 166 30 31 113 21

0.057 0.14 0.18 0.083 0.367 0.34 0.22 0.129 0.168 0.165 0.020 0.21 0.20 1.0 0.32 0.130 0.150 0.052 0.397 0.071 0.074 0.27 0.050

14 20 21 24 86 50 23.5 23.6 24 24.5 3.0 29 53 170 112 23 38 9 62 13 10 41 12

0.137 0.20 0.208 0.236 0.86 0.50 0.235 0.236 0.24 0.245 0.030 0.286 0.53 1.705 1.12 0.23 0.38 0.092 0.62 0.13 0.10 0.41 0.12

TABLE 2 THERMAL DIFFUSIVITIES OF COMMON ALLOYS FROM 20 TO 100 °C (68 TO 212 °F)

ALLOYS

DENSITY g/cm3

ALUMINUM ALLOYS 1100 2.71 2014 2.80 5052 2.68 6061 2.70 7075 2.80 COPPER ALLOYS COMMERCIAL 8.80 BRONZE CARTRIDGE 8.53 BRASS NAVAL BRASS 8.41 BERYLLIUM 8.23 COPPER 9% ALUMINUM 7.58 BRONZE MAGNESIUM ALLOYS AZ 31 1.78 AZ 91 1.83 ZW 1 1.8 RZ 5 1.84 STAINLESS STEELS

HEAT CAPACITY

THERMAL CONDUCTIVITY

THERMAL DIFFUSIVITY

lb/in.3

j/kg · k

calit/g · °c

w/m · k

calit/cm · s · °c

mm2/s

cm2/s

0.098 0.101 0.097 0.098 0.101

963 963 963 963 963

0.23 0.23 0.23 0.23 0.23

222 193 138 172 121

0.53 0.46 0.33 0.41 0.29

85 71 54 66 45

0.85 0.71 0.54 0.66 0.45

0.318 377

0.09

188

0.45

57

0.57

0.308 377

0.09

121

0.29

38

0.38

0.303 377 0.297 419

0.09 0.1

117 84

0.28 0.20

37 24

0.37 0.24

0.273 435

0.104

60

0.144

18

0.18

0.064 0.066 0.065 0.066

0.25 0.24 0.24 0.23

84 84 134 113

0.20 0.20 0.32 0.27

45 46 74 64

0.45 0.46 0.74 0.64

1050 1005 1005 963

TYPE 301 7.9 TYPE 304 7.9 TYPE 316 8.0 TYPE 410 7.7 TYPE 430 7.7 TYPE 501 7.7 NICKEL-BASE ALLOYS NIMONIC 80A 8.19 INCONEL 600 8.42 MONEL 400 8.83 TITANIUM ALLOYS TI-6AL-4V 4.43 TI-5AL-2.5SN 4.46

0.285 0.285 0.289 0.278 0.278 0.278

502 502 502 460 460 460

0.12 0.12 0.12 0.11 0.11 0.11

16 15.1 15.5 24 26 37

0.039 0.036 0.037 0.057 0.062 0.088

4.1 3.8 3.9 6.7 7.3 10

0.041 0.038 0.039 0.067 0.073 0.10

0.296 460 0.304 460 0.319 419

0.11 0.11 0.10

11 15 22

0.027 0.035 0.052

3.0 3.8 5.8

0.030 0.038 0.058

0.160 611 0.161 460

0.146 0.11

5.9 6.3

0.014 0.015

2.1 3.1

0.021 0.031

For the planar heat source on a steel surface, as represented by Fig. 1, the time in seconds to produce melting on the surface, tm, is given by:

TM = (5000/H.I.)2

(EQ 2)

where H.I. is the net heat intensity (in W/cm2) transferred to the workpiece. Equation 2 provides a rough estimate of the time required to produce melting, and is based upon the thermal diffusivity of steel. Materials with higher thermal diffusivities--or the use of a local point heat source rather than a planar heat source-will increase the time to produce melting by a factor of up to two to five times. On the other hand, thin materials tend to heat more quickly. If the time to melting is considered to be a characteristic interaction time, tI, then the graph shown in Fig. 3 can be generated. Heat sources with power densities that are of the order of 1000 W/cm2, such as oxyacetylene flames or electroslag welding, require interaction times of 25 s with steel, whereas laser and electron beams, at 1 MW/cm2, need interaction times on the order of only 25 μs. If this interaction time is divided into the heat-source diameter, dH, then a maximum travel speed, Vmax, is obtained for the welding process (Fig. 4).

FIG. 3 TYPICAL WELD POOL-HEAT SOURCE INTERACTION TIMES AS FUNCTION OF HEAT-SOURCE INTENSITY.

MATERIALS WITH A HIGH THERMAL DIFFUSIVITY, SUCH AS COPPER OR ALUMINUM, WOULD LIE NEAR THE TOP OF THIS BAND, WHEREAS STEELS, NICKEL ALLOYS, OR TITANIUM WOULD LIE IN THE MIDDLE. URANIUM AND CERAMICS, WITH VERY LOW THERMAL DIFFUSIVITIES, WOULD LIE NEAR THE BOTTOM OF THE BAND.

FIG. 4 MAXIMUM WELD TRAVEL VELOCITY AS A FUNCTION OF HEAT-SOURCE INTENSITY BASED ON TYPICAL HEAT-SOURCE SPOT DIAMETERS

The reason why welders begin their training with the oxyacetylene process should be clear: it is inherently slow and does not require rapid response time in order to control the size of the weld puddle. Greater skill is needed to control the morerapid fluctuations in arc processes. The weld pool created by the high-heat-intensity processes, such as laser-beam and electron-beam welding, cannot be humanly controlled and must therefore be automated. This need to automate leads to increased capital costs. On an approximate basis, the W/cm2 of a process can be substituted with the dollar cost of the capital equipment. With reference to Fig. 2, the cost of oxyacetylene welding equipment is nearly $1000, whereas a fully automated laser-beam or electron-beam system can cost $1 million. Note that the capital cost includes only the energy source, control system, fixturing, and materials handling equipment. It does not include operating maintenance or inspection costs, which can vary widely depending on the specific application. For constant total power, a decrease in the spot size will produce a squared increase in the heat intensity. This is one of the reasons why the spot size decreases with increasing heat intensity (Fig. 4). It is easier to make the spot smaller than it is to increase the power rating of the equipment. In addition, only a small volume of material usually needs to be melted. If the spot size were kept constant and the input power were squared in order to obtain higher densities, then the volume of fused metal would increase dramatically, with no beneficial effect. However, a decreasing spot size, coupled with a decreased interaction time at higher power densities, compounds the problem of controlling the higher-heat-intensity process. A shorter interaction time means that the sensors and controllers necessary for automation must operate at higher frequencies. The smaller spot size means that the positioning of the heat source must be even more precise, that is, on the order of the heat-source diameter, dH. The control frequency must be greater than the travel velocity divided by the diameter of the heat source. For processes that operate near the maximum travel velocity, this is the inverse of the process interaction time, tI (Fig. 3). Thus, not only must the high-heat-intensity processes be automated because of an inherently high travel speed, but the fixturing requirements become greater, and the control systems and sensors must have ever-higher frequency responses. These factors lead to increased costs, which is one reason that the very productive laser-beam and electron-beam welding processes have not found wider use. The approximate productivity of selected welding processes, expressed as length of weld produced per second, to the relative capital cost of equipment is shown in Fig. 5.

FIG. 5 APPROXIMATE RELATIONSHIP BETWEEN CAPITAL COST OF WELDING EQUIPMENT AND SPEED AT WHICH SHEET METAL JOINTS CAN BE PRODUCED

Another important welding process parameter that is related to the power density of the heat source is the width of the heat-affected zone (HAZ). This zone is adjacent to the weld metal and is not melted itself but is structurally changed because of the heat of welding. Using the Einstein equation, the HAZ width can be estimated from the process interaction time and the thermal diffusivity of the material. This is shown in Fig. 6, with one slight modification. At levels above approximately 104 W/cm2, the HAZ width becomes roughly constant. This is due to the fact that the HAZ grows during the heating stage at power densities that are below 104 W/cm2, but at higher power densities it grows during the cooling cycle. Thus, at low power densities, the HAZ width is controlled by the interaction time, whereas at high power densities, it is independent of the heat-source interaction time. In the latter case, the HAZ width grows during the cooling cycle as the heat of fusion is removed from the weld metal, and is proportional to the fusion zone width.

FIG. 6 RANGE OF WELD HAZ WIDTHS AS FUNCTION OF HEAT-SOURCE INTENSITY

The change of slope in Fig. 6 also represents the heat intensity at which the heat utilization efficiency of the process changes. At high heat intensities, nearly all of the heat is used to melt the material and little is wasted in preheating the surroundings. As heat intensity decreases, this efficiency is reduced. For arc welding, as little as half of the heat generated may enter the plate, and only 40% of this heat is used to fuse the metal. For oxyacetylene welding, the heat entering the metal may be 10% or less of the total heat, and the heat necessary to fuse the metal may be less than 2% of the total heat. A final point is that the heat intensity also controls the depth-to-width ratio of the molten pool. This value can vary from 0.1 in low-heat-intensity processes to more than 10 in high-heat-intensity processes. It should now be evident that all fusion welding processes can be characterized generally by heat-source intensity. The properties of any new heat source can be estimated readily from the figures in this article. Nonetheless, it is useful to more fully understand each of the common welding heat sources, such as flames, arcs, laser beams, electron beams, and electrical resistance. These are described in separate articles in the Section "Fusion Welding Processes" in this Volume. Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

Introduction DURING FUSION WELDING, the thermal cycles produced by the moving heat source cause physical state changes, metallurgical phase transformation, and transient thermal stress and metal movement. After welding is completed, the finished product may contain physical discontinuities that are due to excessively rapid solidification, or adverse microstructures that are due to inappropriate cooling, or residual stress and distortion that are due to the existence of incompatible plastic strains. In order to analyze these problems, this article presents an analysis of welding heat flow, focusing on the heat flow in the fusion welding process. The primary objective of welding heat flow modeling is to provide a mathematical tool for thermal data analysis, design iterations, or the systematic investigation of the thermal characteristics of any welding parameters. Exact comparisons with experimental measurements may not be feasible, unless some calibration through the experimental verification procedure is conducted. Welding Thermal Process. A physical model of the welding system is shown in Fig. 1. The welding heat source moves at a constant speed along a straight path. The end result, after either initiating or terminating the heat source, is the formation of a transient thermal state in the weldment. At some point after heat-source initiation but before termination, the temperature distribution is stationary, or in thermal equilibrium, with respect to the moving coordinates. The origin of the moving coordinates coincides with the center of the heat source. The intense welding heat melts the metal and forms a molten pool. Some of the heat is conducted into the base metal and some is lost from either the arc column or the metal surface to the environment surrounding the plate. Three metallurgical zones are formed in the plate upon completion of the thermal cycle: the weld-metal (WM) zone, the heated-affected zone (HAZ), and the base-metal (BM) zone. The peak temperature and the subsequent cooling rates determine the HAZ structures, whereas the thermal gradients, the solidification rates, and the cooling rates at the liquid-solid pool boundary determine the solidification structure of the WM zone. The size and flow direction of the pool determines the amount of dilution and weld penetration. The material response in the temperature range near melting temperatures is primarily responsible for the metallurgical changes.

FIG. 1 SCHEMATIC OF THE WELDING THERMAL MODEL

Two thermal states, quasi-stationary and transient, are associated with the welding process. The transient thermal response occurs during the source initiation and termination stages of welding, the latter of which is of greater metallurgical interest. Hot cracking usually begins in the transient zone, because of the nonequilibrium solidification of the base material. A crack that forms in the source-initiation stage may propagate along the weld if the solidification strains sufficiently multiply in the wake of the welding heat source. During source termination, the weld pool solidifies several times faster than the weld metal in the quasi-stationary state. Cracks usually appear in the weld crater and may propagate along the weld. Another dominant transient phenomenon occurs when a short repair weld is made to a weldment. Rapid cooling results in a brittle HAZ structure and either causes cracking problems or creates a site for fatigue-crack initiation. The quasi-stationary thermal state represents a steady thermal response of the weldment in respect to the moving heat source. The majority of the thermal expansion and shrinkage in the base material occurs during the quasi-stationary thermal cycles. Residual stress and weld distortion are the thermal stress and strain that remain in the weldment after completion of the thermal cycle. Relation to Welding Engineering Problems. To model and analyze the thermal process, an understanding of

thermally induced welding problems is important. A simplified modeling scheme, with adequate assumptions for specific problems, is possible for practical applications without using complex mathematical manipulations. The relationship between the thermal behavior of weldments and the metallurgy, control, and distortion associated with welding is summarized below. Welding Metallurgy. As already noted, defective metallurgical structures in the HAZ and cracking in the WM usually

occur under the transient thermal condition. Therefore, a transient thermal model is needed to analyze cracking and embrittlement problems. To evaluate the various welding conditions for process qualification, the quasi-stationary thermal responses of the weld material need to be analyzed. The minimum required amount of welding heat input within the allowable welding speed range must be determined in order to avoid rapid solidification and cooling of the weldment Preheating may be necessary if the proper thermal conditions cannot be obtained under the specified welding procedure. A quasi-stationary thermal model is adequate for this type of analysis. Hot cracking results from the combined effects of strain and metallurgy. The strain effect results from weld-metal displacement at near-melting temperatures, because of solidification shrinkage and weldment restraint. The metallurgical effect relates to the segregation of alloying elements and the formation of the eutectic during the high nonequilibrium solidification process. Using metallurgical theories, it is possible to determine the chemical segregation, the amounts and distributions of the eutectic, the magnitudes and directions of grain growth, and the weld-metal displacement at high temperatures. Using the heating and cooling rates, as well as the retention period predicted by modeling and analysis, hotcracking tendencies can be determined. To analyze these tendencies, it is important to employ a more accurate numerical model that considers finite welding heat distribution, latent heat, and surface heat loss.

Welding Control. In-process welding control has been studied recently. Many of the investigations are aimed at

developing sensing and control hardware. However, a link between weld-pool geometry and weld quality has not been fully established. A transient heat-flow analysis needs to be used to correlate the melted surface, which is considered to be the primary control variable, to the weld thermal response in a time domain. Welding Distortion The temperature history and distortion caused by the welding thermal process creates nonlinear

thermal strains in the weldment. Thermal stresses are induced if any incompatible strains exist in the weld. Plastic strains are formed when the thermal stresses are higher than the material yield stress. Incompatible plastic strains accumulate over the thermal process and result in residual stress and distortion of the final weldment. The material response in the lower temperature range during the cooling cycle is responsible for the residual stresses and weldment distortion. For this type of analysis, the temperature field away from the welding heat source is needed for the modeling of the heating and cooling cycle during and after welding. A quasi-stationary thermal model with a concentrated moving heat source can predict, with reasonable accuracy, the temperature information for the subsequent stress and distortion analysis. Literature Review. Many investigators have analytically, numerically, and experimentally studied welding heat-flow

modeling and analysis (Ref 1, 2, 3, 4, 5, 6, 7, 8, 9, 10, 11, 12, 13, 14, 15, 16, 17, 18). The majority of the studies were concerned with the quasi-stationary thermal state. Lance and Martin (Ref 1), Rosenthal and Schmerber (Ref 2) and Rykalin (Ref 3) independently obtained an analytical temperature solution for the quasi-stationary state using a point or line heat source moving along a straight line on a semi-infinite body. A solution for plates of finite thickness was later obtained by many investigators using the imaged heat source method (Ref 3, 4). Tsai (Ref 5) developed an analytical solution for a model that incorporated a welding heat source with a skewed Gaussian distribution and finite plate thickness. It was later called the "finite source theory" (Ref 6). With the advancement of computer technology and the development of numerical techniques like the finite-difference and finite-element methods, more exact welding thermal models were studied and additional phenomena were considered, including nonlinear thermal properties, finite heat-source distributions, latent heat, and various joint geometries. Tsai (Ref 5), Pavelic (Ref 7), Kou (Ref 8), Kogan (Ref 9), and Brody (Ref 10) studied the simulation of the welding process using the finite-difference scheme. Hibbitt and Marcal (Ref 11), Friedman (Ref 12), and Paley (Ref 13) made some progress in welding simulation using the finite-element method. Analytical solutions for transient welding heat flow in a plate were first studied by Naka (Ref 14), Rykalin (Ref 3), and Masubuchi and Kusuda (Ref 15) in the 1940s and 1950s. A point or line heat source, constant thermal properties, and adiabatic boundary conditions were assumed. Later, Tsai (Ref 16) extended the analytical solution to incorporate Gaussian heat distribution using the principle of superposition. The solution was used to investigate the effect of pulsed conditions on weld-pool formation and solidification without the consideration of latent heat and nonlinear thermal properties. The analysis of the transient thermal behavior of weldments using numerical methods has been the focus of several investigations since 1980. Friedman (Ref 17) discussed the finite-element approach to the general transient thermal analysis of the welding process. Brody (Ref 10) developed a two-dimensional transient heat flow model using a finitedifference scheme and a simulated pulsed-current gas-tungsten arc welding process (GTAW). Tsai and Fan (Ref 18) modeled the two-dimensional transient welding heat flow using a finite-element scheme to study the transient welding thermal behavior of the weldment. General Approach. The various modeling and analysis schemes summarized above can be used to investigate the

thermal process of different welding applications. With adequate assumptions, analytical solutions for the simplified model can be used to analyze welding problems that show a linear response to the heat source if the solutions are properly calibrated by experimental tests. Numerical solutions that incorporate nonlinear thermal characteristics of weldments are usually required for investigating the weld-pool growth or solidification behavior. Numerical solutions can also be necessary for metallurgical studies in the weld HAZ if the rapid cooling phenomenon is significant under an adverse welding environment, such as welding under water. Thermally related welding problems can be categorized as: • • •

SOLIDIFICATION RATES IN THE WELD POOL COOLING RATES IN THE HAZ AND ITS VICINITY THERMAL STRAINS IN THE GENERAL DOMAIN OF THE WELDMENT

The domain of concern in the weld pool solidification is within the molten pool area, in which the arc (or other heat source) phenomena and the liquid stirring effect are significant. A convective heat-transfer model with a moving boundary at the melting temperature is needed to study the first category, and numerical schemes are usually required, as well. The HAZ is always bounded on one side by the liquid-solid interface during welding. This inner-boundary condition is the solidus temperature of the material. The liquid weld pool might be eliminated from thermal modeling if the interface could be identified. A conduction heat-transfer model would be sufficient for the analysis of the HAZ. Numerical methods are often employed and very accurate results can be obtained. The thermal strains caused by welding thermal cycles are caused by the nonlinear temperature distribution in the general domain of the weldment. Because the temperature in the material near the welding heat source is high, very little stress can be accumulated from the thermal strains. This is due to low rigidity, that is, small modulus of elasticity and low yield strength. The domain for thermal strain study is less sensitive to the arc and fluid-flow phenomena and needs only a relatively simple thermal model. Analytical solutions with minor manipulations often provide satisfactory results. In this article, only the analytical heat-flow solutions and their practical applications are addressed. The numerical conduction solutions and the convective models for fluid flow in a molten weld pool are not presented.

References

1. N.S. BOULTON AND H.E. LANCE-MARTIN, RESIDUAL STRESSES IN ARC WELDED PLATES, PROC. INST. MECH. ENG., VOL 33, 1986, P 295 2. D. ROSENTHAL AND R. SCHMERBER, THERMAL STUDY OF ARC WELDING, WELD. J., VOL 17 (NO. 4), 1983, P 2S 3. N.N. RYKALIN, "CALCULATIONS OF THERMAL PROCESSES IN WELDING," MASHGIZ, MOSCOW, 1951 4. K. MASUBUCHI, ANALYSIS OF WELDED STRUCTURES, PERGAMON PRESS, 1980 5. C.L. TSAI, "PARAMETRIC STUDY ON COOLING PHENOMENA IN UNDERWATER WELDING," PH.D THESIS, MIT, 1977 6. C.L. TSAI, FINITE SOURCE THEORY, MODELING OF CASTING AND WELDING PROCESSES II, ENGINEERING FOUNDATION MEETING, NEW ENGLAND COLLEGE (HENNIKER, NH), 31 JULY TO 5 AUG 1983, P 329 7. R. PAVELIC, R. TANAKUCHI, O. CZEHARA, AND P. MYERS, EXPERIMENTAL AND COMPUTED TEMPERATURE HISTORIES IN GAS TUNGSTEN ARC WELDING IN THIN PLATES, WELD. J., VOL 48 (NO. 7), 1969, P 295S 8. S. KOU, 3-DIMENSIONAL HEAT FLOW DURING FUSION WELDING, PROC. OF METALLURGICAL SOCIETY OF AIME, AUG 1980, P 129-138 9. P.G. KOGAN, THE TEMPERATURE FIELD IN THE WELD ZONE, AVE. SVARKA, VOL 4 (NO. 9), 1979, P 8 10. G.M. ECER, H.D. DOWNS, H.D. BRODY, AND M.A. GOKHALE, HEAT FLOW SIMULATION OF PULSED CURRENT GAS TUNGSTEN ARC WELDING, MODELING OF CASTING AND WELDING PROCESSES, ENGINEERING FOUNDATION 1980 MEETING (RINDGE, NH), 3-8 AUG 1980, P 139160 11. H. HIBBITT AND P. MARCAL, A NUMERICAL THERMOMECHANICAL MODEL FOR WELDING AND SUBSEQUENT LOADING OF A FABRICATED STRUCTURE, COMPUT. STRUCT., VOL 3, 1973, P 1145 12. E. FRIEDMAN, THERMOMECHANICAL ANALYSIS OF THE WELDING PROCESS USING FINITE ELEMENT METHODS, TRANS. ASME, AUG 1975, P 206 13. Z. PALEY AND P. HIBBERT, COMPUTATION OF TEMPERATURE IN ACTUAL WELD DESIGN, WELD. J., VOL 54 (NO. 11), 1975, P 385.S

14. T. NAKA, TEMPERATURE DISTRIBUTION DURING WELDING, J. JPN. WELD. SOC., VOL 11 (NO. 1), 1941, P 4 15. K. MASUBUCHI AND T. KUSUDA, TEMPERATURE DISTRIBUTION OF WELDED PLATES, J. JPN WELD. SOC., VOL 22 (NO. 5), 1953, P 14 16. C.L. TSAI AND C.A. HOU, THEORETICAL ANALYSIS OF WELD POLL BEHAVIOR IN THE PULSED CURRENT GTAW PROCESS, TRANSPORT PHENOMENA IN MATERIALS PROCESSING, ASME WINTER ANNUAL MEETING, 1983 17. E. FRIEDMAN, "FINITE ELEMENT ANALYSIS OF ARC WELDING," REPORT WAPD-TM-1438, DEPARTMENT OF ENERGY, 1980 18. J.S. FAN AND C.L. TSAI, "FINITE ELEMENT ANALYSIS OF WELDING THERMAL BEHAVIOR IN TRANSIENT CONDITIONS," 84-HT-80, ASME Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

Mathematical Formulations Conduction Equation. A diagram of the welding thermal model is shown in Fig. 1. The origin of the moving coordinates (w,x,z) is fixed at the center of the welding heat source. The coordinates move with the source at the same speed. The conduction equation for heat flow in the weldments is:

∇.(λ∇θ ) + ρ C p v

∂θ • ∂θ + Q = ρC pv ∂w ∂t

(EQ 1)

The initial condition is:

θ= θ 0, AT T = 0

(EQ 2)

and the general boundary condition is: λ

• ∂θ ∂θ ∂θ Iw + Iy + I z − q + h(θ − θ ∞ ) = 0 ∂w ∂y ∂z

(EQ 3)

where ∇ is a differential operator; θ is the temperature; θ ∞ is the environmental temperature; θ0 is the initial temperature; λ is thermal conductivity; ρ is density; Cp is specific heat h is the surface heat-loss coefficient; lw, ly, and lz are the direction cosines of the boundary surface; Q is the volumetric heat source, t is time, and v is welding speed. The volumetric heat source represents the Joule heating in the weldment that is due to the electric current flow within that conducting medium. The total energy of such heating in welding is usually minimal, compared to the arc heat input. The majority of the energy is concentrated in a very small volume beneath the arc (Ref 5). In other words, a very high energy density generation exists in the weld pool, and it may have a significant effect on transient pool growth and solidification. Heat-Source Formulation. The direction cosines on the surface that receive the heat flux from the welding source (z =

0) are lw = ly = 0 and lz = -1. Within the significant heat-input area (to be defined later in this section), the heat loss coefficient, h, is zero. The distribution of the welding heat flux on the weldment surface can be characterized, in a general form, by a skewed Gaussian function (Ref 19): • • βv   q (r , w) = q 0 exp  −Cr ² − w 2k  

(EQ 4)

•

where β is a weight constant, κ is the thermal diffusivity of the base material, C is a shape constant, q is heat flux as a •

function of (r, w), q 0 is heat flux at the source center, r is the radial coordinate from the source center, and v is welding speed. The weight constant, β, indicates the significance of the welding travel speed. A normal distribution of the welding heat flux is obtained if the weight constant is zero. In general, the total energy input to the weldment, which is a fraction of the total welding power generated by the welding machine, is the sum of the concentrated heat and the diffused heat (Ref 20). The concentrated heat is carried by the core of the energy transmission medium, for example, the arc plasma column. The diffused heat reaches the weld surface by radiation and convection energy transport from the core surface. The heat-flux distribution is a function of the proportional values between these two types of energy. The fraction of the total welding power reaching the weldment indicates the heating efficiency of the welding process, and the fraction percentage is defined as welding heat efficiency, η. The shape constant, C, can be obtained in terms of the core diameter, D, and the concentration factor, F. The concentration factor is defined as the ratio of the concentrated heat to the net energy reaching the weldment. The core diameter can be assumed to be the diameter of the plasma column in the arc welding process. The concentration factor and welding heat efficiency are not fully understood and have been subjected to manipulation during the mathematical analyses in order to obtain a better correlation with the experimental data. Assuming a normal heat-flux model, two concentrations are required to determine the shape constant and the heat flux at •

the source center, q 0. By integrating Eq 4 over the core heat area and the entire heat input domain (r = 0 → ∞ ), the shape factor can be determined by dividing the two integrals. The heat flux at the source center can then be determined from the second integral. The two constants are expressed as: C=

4larc [1/(1 − F )] D²

(EQ 5)

where larc is the length of the arc plasma column. •

q0 •

=

Q

C π

(EQ 6)

In the case of arc welding, •

q0 C = η EI π

(EQ 7)

where E is the welding arc voltage and I is the welding current. For practical purposes, the welding heat source can be considered to be restricted within a circle of radius ra, where the •

heat flux drops to 1/100 of the center flux q 0. The radius of the significant heat input area can be written as:  I 100  ra =  arc   C 

0.5

(EQ 8)

Surface Heat Loss. The heat-loss coefficient, h, represents both radiation and convection heat loss from the boundary surfaces outside the significant heat input area. The formulation for both heat-loss mechanisms can be written as the radiation heat-loss coefficient (in air):

HRAD = εσ(θw +

∞ )(

θW 2- θ ∞ 2))

(EQ 9)

or the natural convection heat-loss coefficient (in air): H AC = 0.00042

θW − θ ∞ B

(EQ 10)

or the convection heat-loss coefficient (in water):

HCW = 0.442(θW - θ ∞ )0.25

(EQ 11)

where ε is emissivity, σ is the Stefan-Boltzmann constant, θw is the surface temperature, θ ∞ , is the environmental temperature, and B is the characteristic surface dimension. Natural convection is dominant at a temperature below 550 °C (1020 °F), whereas radiation becomes more important at temperatures above this level. The total heat-loss coefficient is the sum of Eq 9 and 10. The characteristic surface dimension is the effective distance from the source beyond which the temperature rises insignificantly during welding. The characteristic dimension for steel is about 150 mm (6 in.) (Ref 5). In underwater welding, heat losses are primarily due to heat transfer from the surface to the moving water environment. This motion is created by the rising gas column in the arc area (Ref 21). For an insulated surface, no heat transfer into or out of the surface is assumed. The temperature gradient normal to the surface is zero, and can be represented by:

N· ∇θ=0

(EQ 12)

where n is a unit vector normal to the surface and equals (lw2 + ly2 + lz2)0.5. Other Boundary Conditions. There are several other possible boundary conditions in welding heat-flow modeling that

depend on the assumptions used for model simplification. One is the condition at infinity: θ =θ ∞ or lim r →∞

∂θ =0 ∂r

(EQ 13)

Another is the condition near the heat source. In the case of a line source for a thin plate: -2πλ H lim r r →∞

∂θ = η EI ∂r

(EQ 14)

In the case of a point source for a thick plate: -2πλ lim r ² r →∞

∂θ = η EI ∂r

(EQ 15)

In the case of a finite source for a thick plate: •

N · (-λ ∇ θ) = q ; R ≤ RA Another is represented by the conditions at the solid-liquid interface:

(EQ 16)

θ1 = θS = θM n.(λ∇θ )s − n(λ∇θ )1 = ± ρ s L

(EQ 17) ds dt

(EQ 18)

where + indicates the melting process and - indicates the solidification process. The subscripts s and l indicate the temperature and the properties in a solid and liquid, respectively. The n is a normal vector on the boundary surface or interface, ra is the radius of the heat-input area, L is the latent heat of the base material, and the subscript m represents the melting temperature of the base material.

References cited in this section

5. C.L. TSAI, "PARAMETRIC STUDY ON COOLING PHENOMENA IN UNDERWATER WELDING," PH.D THESIS, MIT, 1977 19. R.L. APPS AND D.R. MILNER, HEAT FLOW IN ARGON-ARC WELDING, BR. WELD. J., VOL 2 (NO. 10), 1955, P 475 20. H.S. CARSLAW AND J.C. JAEGER, CONDUCTION OF HEAT IN SOLIDS, OXFORD PRESS 21. C.L. TSAI AND J.H. WU, "AN INVESTIGATION OF HEAT TRANSPORT PHENOMENA IN UNDERWATER WELDING," PRESENTED AT THE ASME WINTER ANNUAL MEETING (MIAMI BEACH, FL), 1985

Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

Engineering Solutions and Empirical Correlation General Solutions. The general (analytical) heat-flow solutions for fusion welding can be categorized by those appropriate for a thick plate, a thin plate, or a plate with finite thickness. In most cases, the boundary surfaces (except for the heat-input area) are assumed to be adiabatic, and the thermal properties are independent of temperature. The various metallurgical zones in the weldment are assumed to be homogeneous, and the thermal model is linear.

The solutions give the temperature for a specific point if the welding velocity, v, voltage, E, and current, I, as well as the physical properties of the plate material (ρ, λ, Cp) and the welding heat efficiency, η, are known. This specific point is defined by r and w in: r = w² + y ² + z ²

(EQ 19)

where w = x - vt. The heat-flow solutions are not accurate at points near the welding arc, because a point source or line source is assumed for thick and thin plates, respectively. To approximate the transient temperature changes at the start and end of a weld, Fig. 2 shows a global coordinate system (x,y,z), the origin of which is fixed at the source initiation, where t0 is the welding time and t1 is the time after the welding heat-source termination. The temperature solutions at t0 and t1 are the temperature changes at the start and end of the weld, respectively.

FIG. 2 GLOBAL AND MOVING COORDINATE SYSTEMS FOR WELDING HEAT CONDUCTION.

The temperature solution for thick plate at the arc start location is: η EI 2πλ vto

θ − θ0 =

(EQ 20)

The quasi-stationary temperature distribution is: θ − θ0 =

η EI  −v ( w + r )  exp   2πλ r  2κ 

(EQ 21)

At the arc termination location, the solution is: θ − θ0 =

η EI 2πλ vt1

(EQ 22)

The temperature solution for thin plate at the arc start location is:

θ − θ0 =

v ²t v ²t η EI exp o ko o 2πλ H 2κ 2κ

(EQ 23)

The quasi-stationary temperature distribution is: θ − θ0 =

η EI −vw vr exp ko 2πλ H 2κ 2κ

(EQ 24)

For the arc termination location, the solution is: θ − θ0 =

v ²t η EI  v ²t  exp 1 ko  1  2πλ H 2κ  2κ 

(EQ 25)

where K0 is the modified Bessel function of the second kind of zeroth order and η EI is the welding heat input rate. Temperature for Plate With Finite Thickness. The image method enables the investigator to superimpose the

solutions for an infinitely thick plate, the source of which is placed on imaginary surfaces until the proper boundary

conditions on the plate surfaces are obtained. This method is based on the premise that if a solution satisfies the governing equation and the boundary conditions, then it must be not only a correct solution, but the only solution (that is, the uniqueness of solution premise). Using the image method, the solution for plates of finite thickness with a adiabatic surfaces can be modified from the respective temperature solutions described previously. Let φ 0 (w,y,z,t) be the initial solution for an infinitely thick plate. The temperature solution for a finite thick plate can be obtained by super imposing the imaginary solutions φ mn (w,ym, zn, t) and φ 'mn (w,y' m, z'n, t) to the initial solution, and this can be written in a general form as: ∞

∞

θ − θ 0 = −φ0 ( w, y, z , f ) + ∑ ∑ [φmn ( w, ym , zm , t ] + φ 'mn ( w, y 'm , Z m t )]

(EQ 26)

m=0 n=0

where ym = 2mB - y; y'm = 2mB + y; zn = 2nH - z; and z'n = 2nH + z, in which B is the half width and H is the thickness of the plate. The subscripts m and n are integers that vary from zero to infinity. For a plate with sufficient width, the subscript m is zero. The solution will coverage and reach the correct adiabatic surface condition in six to ten superposition steps, depending on the thickness of the plate. The two-dimensional solution (that is, thin plate) is generally used for any solution that requires more than ten superposition steps. Equation 26 can be expressed as:   −vr2 n  −vw    −vr  η EI exp    exp   ∞  exp  2k 2 k 2 k     +   θ − θ0 = −  ∑ 2πλ r r2 n  n = 1    

   −vr2 n +1     exp  2k    +    r2 n −1     

(EQ 27)

where

R2N = W2 + Y2 + (2NH - Z)2 AND R2N+1 = W2 + Y2 + (2NH + Z)2

(EQ 28)

Cooling Rate. Frequently, it is desirable to know the cooling rate experienced at some location in a weldment to enable a

prediction of the metallurgical structure in that area. A general methodology by which cooling rate equations are obtained from the temperature-distribution equations is discussed below. Recall that the moving coordinate w is defined by w = x - vt. Using this definition, it is easily shown that: ∂w = −v ∂t

(EQ 29)

Using the chain rule, the cooling-rate equation is: ∂θ ∂θ = −v ∂t ∂w

(EQ 30)

Because the temperature-distribution equations are a function of w and r, the cooling-rate equations can be obtained by differentiating the temperature-distribution equations with respect to w and multiplying by -v. The cooling rate is defined as the slope of a tangent line drawn on the temperature-time curve. Because the cooling rate changes with temperature, when one speaks of a cooling rate, the specific temperature, θc, at which it occurs must also be

given. In a weldment, the variable of interest is the cooling rate at the critical temperature that ultimately defines what type of metallurgical structure will result (if the material is heat treatable). For steels, this critical temperature is the "nose" of the continuous cooling-transformation (CCT) curve. At this temperature, the cooling rate determines if upper transformation products (pearlite, upper bainite) or lower transformation products (martensite, lower bainite) will form. For many steels, this critical temperature ranges from approximately 200 to 540 °C (400 to 1000 °F). The cooling rate in a weldment is also a function of location. In order to find a cooling-rate equation, the particular location in the weldment that is of interest must be defined. The resulting cooling-rate equation will be applicable only to that location. The differentiation, ∂ θ/ ∂ w, of either Eq 21 or 27, which is required to obtain the cooling-rate expression, will result in a function of w and r. The variable r can be written in terms of w if the location of interest is defined by a given set of values of y and z. This relationship for r, once formulated, can then be substituted into ∂ θ/ ∂ w, the result being a function of w alone. To determine w corresponding to the critical temperature, θc, a temperature-distribution equation is required (Eq 21 or 27). The aforementioned r-w relationship and temperature distribution equation (Eq 21 or 27) where θ is equal to θc, critical temperature, are used to determine w. Then w is substituted into the dθ/dw expression obtained previously. The end result will be an equation that defines the cooling rate for a particular location in the weldment, and, being a function of the critical temperature, the welding conditions and thermal conductivity of the base material. To determine the cooling rate in a thick plate along the weld centerline (that is y = 0) for a particular critical temperature, the cooling-rate equation can be reduced to: ∂θ 2πλ v (θ c − θ 0 ) = ∂t η EI

2

(EQ 31) •

This equation has been used to predict weld cooling rates in shop practices. Cooling rate is inversely (that is, Q /V) and is proportional to thermal conductivity and the critical temperature at which the cooling rate needs to be evaluated. On the basis of experimental results, a cooling-rate equation was developed for the HAZ of low-carbon steel weldments (Ref 25). This equation considers the combined effects of plate thickness, H, preheating temperature, θ0, and welding conditions, and is given as: ∂θ ∂t

HAZ

 θ − θ 0 1.7  2 −1  H − H o = 0.35  c 1 + tan    α  I / v   π

       

0.8

(EQ 32)

The variables α and H0 depend on the critical temperature of interest. Several values are given in Table 1.

TABLE 1 SELECTED CRITICAL TEMPERATURE AND CORRESPONDING VALUES FOR α AND H0

CRITICAL TEMPERATURE, θC °C °F 700 1290 540 1000 300 570

α

H0 mm 9.9 14.2 19.8

in. 0.39 0.56 0.78

mm 2.0 4.1 9.9

in. 0.08 0.16 0.39

The units used in Eq 32 are important, because the same units that were used in developing the equation must be employed in its application. The plate thickness, H, must be given in inches, and the travel speed, v, must be given in in./min. The temperatures θc and θ0 must be given in °C, and the welding current, I, must be given in amperes. Using the

correct units, the application of Eq 32 will result in a predicted cooling rate (°C/s) for the HAZ of a low-carbon steel weldment. For low-carbon steels welded by the shielded metal arc welding (SMAW), gas-metal arc welding (GMAW), and submerged arc welding (SAW) processes, an empirical equation has been developed that correlates the weld-metal cooling rate at 538 °C (1000 °F) with a 95 to 150 °C (200 to 300 °F) preheat, the weld nugget area (Ref 26): 1.119

  2012 °C / s =    nugget area 

(EQ 33)

where the nugget area is in mm2. For low-carbon steels, an empirical chart for determining nugget area for a given welding condition (Fig. 3) also has been developed (Ref 27). The straight line drawn to connect the current and travel speed intersects the nugget area at the calculated value. Welding voltage controls weld bead shape.

FIG. 3 RELATION BETWEEN NUGGET AREA, HEAT INPUT, AND CURRENT

Peak Temperature. An equation to determine the peak-temperature in a weldment at a given distance y from the weld

centerline would enable the prediction of HAZ sizes, as well as weld bead widths. The general concept of obtaining a peak-temperature equation, as well as some results that have been obtained, are discussed below. Consider Fig. 4 and note that the maximum, or peak, temperature is given when ∂ θ/ ∂ t = 0. For the thick-plate model, the cooling rate can be obtained by differentiating Eq 21 and multiplying by -v: ∂θ ∂θ −vη EI  −v ( w + r )   − w v  w   = −v = exp   x  r ² − 2k  1 + r   ∂t ∂w 2πλ r  2k    

Clearly, the only way that ∂ θ/ ∂ t can be equal to zero is if

(EQ 34)

−w v  w  − 1 + = 0  r ² 2k  r 

(EQ 35)

FIG. 4 SCHEMATIC SHOWING PEAK TEMPERATURE AT (WP, YP, ZP) WITH WP TO BE DETERMINED FOR A GIVEN PEAK TEMPERATURE VALUE FOR A GIVEN (YP, ZP) LOCATION. (A) ISOTHERMS. (B) TEMPERATURE HISTORY

Equation 35 describes the relationship that must exist between the two location variables, r and w, for the temperature at the point to be equal to the peak temperature. If this expression were to be substituted into the temperature distribution equation for thick plates (Eq 21) and solved for w and r (two equations and two unknowns), then the location of the peak temperature could be determined in terms of w and r. The location given by r and w would be easily converted to y and z as:

R2 = W2 + Y2 + Z2

(EQ 36)

Such a solution for r and w is not explicitly possible, however, because the equations for r and w that result are not explicit. Consequently, iterative techniques are required, resulting in a solution that is both cumbersome and timeconsuming. One method of obtaining a simpler thick-plate peak-temperature equation is to assume that the heat input is from an instantaneous line on the surface of the plate, rather than from a moving point source (that is, v → ∞ ). This allows the elimination of the time dependency in the peak-temperature evaluation. Using this assumption, the temperature distribution is given by: θ − θ0 =

η EI  r²  exp   2πλ t  4κ t 

(EQ 37)

Again, to find the peak-temperature location, ∂ θ/ ∂ t is set equal to zero and the equation is solved for r. The result is: r² =1 4κ t

(EQ 38)

Substituting Eq 38 into Eq 37 yields the peak-temperature expression:  e ρ C pπ r ²   η EI  1 = /  θ − θ0  2   v 

(EQ 39)

It has been found that Eq 39 gives results that are too high, but that the slope of 1/(θp -θ0) versus r2 is accurate. To rectify this situation, Eq 39 is forced to fit experimental results by specifying a known temperature/location condition. When this is done, Eq 39 becomes:    [e ρ C π (r ² − r 2 )] / 2  1 1 p t + = EI η θ p − θ0     θr − θ0      v   

(EQ 40)

where θr and rr are the reference temperature and distance. If the peak temperature (θp) evaluation is restricted to locations on the plate surface (z = 0), and if the reference temperature and distance are assumed to be the melting temperature and the distance from the weld centerline to the fusion boundary (one half of the weld bead width), then Eq 40 can be written as: 2  d   e ρ C pπ ( y ² −   2  1  2 = θ p − θ0   η EI      v    

   1   + θ −θ  m 0   

(EQ 41)

where θm is the melting temperature and d is the weld bead width. This equation gives the peak temperature θp in a thick plate at a distance y from the weld centerline. Solidification Rate. The weld solidification structure can be determined by using the constitutional supercooling

criterion. Three thermal parameters that influence the solidification structure are temperature gradient normal to the solidliquid interface, G (°C/cm), solidification rate of the interface, R (cm/s), and cooling rate at the interface, dθ/dt at melting temperature (°C/s and equal to the product of GR). The microstructure may change from being planar to being cellular, a columnar dendrite, or an equiaxial structure if the G/R ratio becomes smaller. The dendrite arm spacing will decrease as the cooling rate increases. The solidification structure becomes refined at higher cooling rates.

At a quasi-steady state, the weld pool solidifies at a rate that is equal to the component of the electrode travel speed normal to the solid-liquid interface. Therefore, the solidification rate varies along the solid-liquid interface from the electrode travel speed, at the weld trailing edge, to zero, at the maximum pool width. The temperature gradient and the cooling rate at the solid-liquid interface can be determined from Eq 21, 24, and 27. Modified Temperature Solution. The temperature solutions have a singularity at the center of the heat source. This singularity causes the predicted temperatures to be inaccurate in the area surrounding the heat source. However, a condition exists in which the peak temperature along the weld bead edge, that is, the solid-liquid interface location at the maximum pool width, is the melting temperature of the material. Using this temperature condition as a boundary condition for the temperature solutions, Eq 21 and 24 can be modified as shown below for thin plate:  −vw   −vr  θ − θ 0 = Bz exp   K0    2κ   2κ 

(EQ 42)

where Bz is a heat input constant to be determined from the weld bead width, d.  K (vr / 2κ )  θ m − θ 0 exp(vrB / 2κ )  0 B   K1 (vrB / 2κ )  Bz = K 0 (vrB / 2κ )

(EQ 43)

with a bead width of:   K (vr / 2κ )  2  d = 2rB 1 −  0 B   K (vr / 2κ )     1 B 

0.5

(EQ 44)

and for thick plate:  −v ( w + r )  θ m − θ 0 = Bz exp  /r  2κ 

(EQ 45)

 1  −vrB   − 1 Bz = (θ m − θ 0 )rB exp     2κ  1 + (2κ / vrB ) 

(EQ 46)

with a bead width of:   1 d = 2rB 1 −   1 + (2κ / vrB ) 

0.5

(EQ 47)

The welding heat input, Q, is replaced by the weld bead width. A Practical Application of Heat Flow Equations (Ref 22). The thermal condition in and near the weld metal must be

established to control the metallurgical events in welding. The particular items of interests are the: • • •

DISTRIBUTION OF PEAK TEMPERATURE IN THE HAZ COOLING RATES IN THE WELD METAL AND IN THE HAZ SOLIDIFICATION RATE OF THE WELD METAL

Although the following discussion primarily focuses on manual are welding, certain general statements are applicable to all welding processes.

Peak Temperatures. The distribution of peak temperatures in the base metal adjacent to the weld is given by (Ref 23): 4.13ρ C p tY 1 1 = + Tp − T0 H met Tm − T0

(EQ 48)

where Tp is the peak temperature (°C) at distance Y (mm) from the weld fusion boundary, T0 is the initial temperature (°C), Tm is the melting temperature (°C), Hnet is the net energy input equal to ηEI/v (J/s · mm), ρ is the density of the material (g/mm3), Cp is the specific heat of solid metal (J/g · °C), and t is the thickness of the base metal (mm). Equation 48 can be used in order to determine the: • • •

PEAK TEMPERATURES AT SPECIFIC LOCATIONS IN THE HAZ WIDTH OF THE HAZ EFFECT OF PREHEAT ON THE WIDTH OF THE HAZ

In addition, determination of the peak temperature at specific locations in the HAZ and the width of the HAZ can be obtained by the procedure described from Eq 34, 35, 36, 37, 38, 39, 40, and 41. Cooling Rate. Because the cooling rate varies with position and time, its calculation requires the careful specification of

conditions. The most useful method is to determine the cooling rate on the weld centerline at the instant when the metal passes through a particular temperature of interest, Tc. At a temperature well below melting, the cooling rate in the weld and in its immediate HAZ is substantially independent of position. For carbon and low-alloy steels, Tc is the temperature near the pearlite "nose" temperature on the time-temperature transformation (TTT) diagram. The value of Tc = 550 °C (1020 °F) is satisfactory for most steels, although not critical. The cooling rate for thick plate (Ref 23) is: R=

2πλ (Tc − T0 )² H met

(EQ 49)

Equation 49 is comparable to Eq 31, which was obtained by the procedure described in the section "Cooling Rate" in this article. Adams (Ref 24) has developed a cooling rate equation for thin plate along the centerline from Eq 24. The cooling rate for thin plate (Ref 24) is: 2

 t  R = 2πλρ C p   + (Tc − T0 )³  H met 

(EQ 50)

where R is the cooling rate (°C/s) at a point on the weld centerline at just that moment when the point is cooling past the Tc, and λ is the thermal conductivity of the metal (J/mm · s · °C). The dimensionless quantity τ , called the "relative plate thickness," can be used to determine whether the plate is thick or thin:

τ =t

ρ C p (Tc − T0 ) H met

(EQ 51)

The thick-plate equation applies when τ is greater than 0.75, and the thin-plate equation applies when τ is less than that value.

Equations 49 and 50 are used to determine the cooling rate along the centerline for thick plate and thin plate, respectively. If one is interested in the cooling rate at the location at distance y (in mm) from the centerline, iterative techniques should be used to solve the cooling rate. First, w and r can be obtained by iteration of the simultaneous equation, which consists of Eq 21 or 27, where θ equals θc and r2 = w2 + y2, where y is given. Then substitute w and r into the differentiation, ∂ θ/ ∂ t = -v ∂ θ/ ∂ w, from the temperature from Eq 21 or 27. The result will be the cooling rate for thick or thin plate located at y distance from the centerline: ∂θ ∂t θ c

In addition, the cooling rate for HAZ of low-carbon steel weldments can be obtained from Eq 32 directly. The solidification rate can have a significant effect on metallurgical structure, properties, response to heat treatment, and soundness. The solidification time, St, of weld metal, measured in seconds, is: St =

LH net 2πλ C p (Tm − To )²

(EQ 52)

where L is the heat of fusion (J/mm3). Example 1: Welding of 5 mm (0.2 in.) Thick Low-Carbon Steels. The thermal properties needed for heat flow analysis are assumed to be:

MELTING TEMPERATURE (TM), °C (°F) AUSTENIZATION TEMPERATURE, °C (°F) THERMAL CONDUCTIVITY (λ), W/M · K (J/MM · S · °C) VOLUMETRIC SPECIFIC HEAT (ρCP), J/MM3 · °C HEAT OF FUSION (L), J/MM3

1510 (2750) 730 (1350) 11.7 (0.028) 0.0044 2

The welding condition is assumed to be:

CURRENT (I), A ARC VOLTAGE (E), V TRAVEL SPEED (V), MM/S (IN./S) PREHEAT (T0), °C (°F) HEAT-TRANSFER EFFICIENCY (η) NET ENERGY INPUT, HNET, J/MM (KJ/IN.)

200 20 5 (0.2) 25 (77) 0.9 720 (18.3)

Calculation of the HAZ Width. The value of Y at Tp = 730 °C (1345 °F) must be determined from Eq 48: 4.13(0.0044)5Yz 1 1 = 730 − 25 720 1510 − 25

resulting in a value for Yz (the width of the HAZ) of 5.9 mm (0.24 in.).

(EQ 53)

In addition, the HAZ width can be obtained by the procedure described in Eq 34, 35, 36, 37, 38, 39, 40, and 41. By substituting w = wp, r = rp, y = yp, and z = 0 (on surface) into Eq 35 and 36, we obtain:

− wp rp2

−

v  wp 1 + 2κ  rp

  = 0 

rp2 = w2p + y 2p

(EQ 54) (EQ 55)

The substitution of Eq 55 in Eq 54 yields:

− wp w +y 2 p

2 p

=

wp v 1+ 2κ w2p + y 2p

(EQ 56)

Inserting v = 5 mm/s, κ= λ/ρcp = 0.028/0.004 into Eq 56:

− wp w +y 2 p

2 p

= 0.3831 +

wp w2p + y 2p

(EQ 57)

wp and yp should satisfy Eq 57. From the temperature distribution equation (Eq 21) and E = 10V, I = 200 A, η= 0.9, λ= 0.028, θp = 730 °C, θ0 = 25 °C, v = 5 mm/s, κ= λ/ρcp = 0.028/0.044:

730 − 25 =

 −5( wp + rp )  0.9 x 20 x 200 / 5 x exp   2π x0.028rp  2 x0.028 / 0.0044 

-1.76 + LNRP = -0.393 (WP + RP)

(EQ 58) (EQ 59)

Solving Eq 21 using the above values, by iteration: •

•

•

ASSUME RP = 3.4 AND PUT THIS VALUE OF RP INTO EQ 59 TO OBTAIN WP = -2.036. SUBSTITUTE RP = 3.4 AND WP = -2.036 INTO EQ 55 AND SOLVE FOR YP: YP = 2.7. SUBSTITUTE YP AND WP INTO EQ 56: THE LEFT SIDE OF EQ 56 = 0.176; THE RIGHT SIDE OF EQ 56 = 0.158. THE RESULTS DO NOT SATISFY EQ 56. ASSUME RP = 3.5 AND PUT THIS VALUE OF RP INTO EQ 59 TO OBTAIN WP = -2.2. SUBSTITUTE RP = 3.5 AND WP = -2.2 INTO EQ 55 AND SOLVE FOR YP: YP = 2.72. SUBSTITUTE YP AND WP INTO EQ 56: THE LEFT SIDE OF EQ 56 = 0.179; THE RIGHT SIDE OF EQ 56 = 0.146. THE RESULTS DO NOT SATISFY EQ 56. ASSUME RP = 3.3 AND PUT THIS VALUE OF RP INTO EQ 59 TO OBTAIN WP = -1.86. SUBSTITUTE RP = 3.3 AND WP = -1.86 INTO EQ 55 AND SOLVE FOR YP: YP = 2.726. SUBSTITUTE YP AND WP INTO EQ 56: THE LEFT SIDE OF EQ 56 = 0.170; THE RIGHT SIDE OF EQ 56 = 0.171. EQUATION 56 IS NOW SATISFIED.

The full HAZ width equals 2yp = 2 × 2.726 = 5.5 mm. Comparing this result, 5.5 mm, with the result obtained with Eq 48 (that is, 5.9 mm), we know that Eq 48 can be used to obtain accurate results when calculating HAZ width and peak temperature.

Effect of Tempering Temperature on Quenched and Tempered (Q&T) Steels. If the plate had been quenched

and then tempered to 430 °C (810 °F), then any region heated above that temperature will have been "over-tempered" and may exhibit modified properties. It would then be reasonable to consider the modified zone as being "heat affected," with its outer extremity located where Tp = 430 °C (810 °F):

4.13(0.0044)5Yz 1 1 = + 430 − 25 720 1510 − 25

(EQ 60)

resulting in a value for Yz of 14.2 mm (0.568 in.). Effect of Preheating Temperature on Q&T Steels. Assume that the Q&T steel described above was preheated to a

temperature, T0, of 200 °C (390 °F):

4.13(0.0044)5Yz 1 1 = + 430 − 200 720 1510 − 200

(EQ 61)

resulting in a value for Yz of 28.4 mm > 14.2 mm (1.14 > 0.568 in.). Therefore, increasing the preheating temperature will increase the value of Yz. Effect of Energy Input on Q&T Steels. Assume that the energy input into the Q&T steel (without preheating)

increases 50% (that is, 1.08 kJ/mm, or 27.4 kJ/in.):

4.13(0.0044)5Yz 1 1 = + 430 − 200 1080 1510 − 25

(EQ 62)

resulting in a value for Yz of 21.3 mm > 14.2 mm (0.839 > 0.568 in.). Therefore, increasing the energy input will increase the value of Yz. Example 2: Welding of 6 mm (0.24 in.) Thick Low-Carbon Steels. The thermal properties needed for heat-flow analysis are assumed to be:

MELTING TEMPERATURE (TM), °C (°F) AUSTENIZATION TEMPERATURE, °C (°F) THERMAL CONDUCTIVITY ( λ ), W/M · K (J/MM · S · °C) VOLUMETRIC SPECIFIC HEAT ( ρ CP), J/MM3 · °C HEAT OF FUSION (L), J/MM3 The welding condition is assumed to be:

CURRENT (I), A ARC VOLTAGE (E), V PREHEAT (T0), °C (°F) HEAT-TRANSFER EFFICIENCY (η)

300 25 25 (77) 0.9

1510 (2750) 730 (1346) 11.7 (0.028) 0.0044 2

Critical Cooling Rate at 550 °C (1020 °F) (Tc). A critical cooling rate exists for each steel composition. If the actual

cooling rate in the weld metal exceeds this critical value, then hard martensitic structures may develop in the HAZ, and there is a great risk of cracking under the influence of thermal stresses in the presence of hydrogen. The best way to determine the critical cooling rate is to make a series of bead-on-plate weld passes in which all parameters, except the arc travel speed, are held constant. After the hardness tests on the weld passes deposited at travel speeds of 6, 7, 8, 9 and 10 mm/s (0.23, 0.28, 0.32, 0.35, and 0.39 in./s), it was found that at the latter two travel speeds, the weld HAZ had the highest hardness. Therefore, the critical cooling rate was encountered at a travel speed of approximately 8 mm/s (0.32 in.s). At this speed, the net energy input is:

H net =

25(300)0.9 = 8.43.75 J / mm 8

(EQ 63)

From Eq 51, the relative plate thickness is:

τ =6

0.0044(550 − 25) = 0.31 843.75

(EQ 64)

Because τ is less than 0.75, the thin-plate equation (Eq 50) applies: 2

R 3  6  = 0.0044   ( 550 − 25 ) = 32.2 2πλ  843.75 

(EQ 65)

resulting in R being equal to 2π(0.028)32.2, which is equal to 5.7 °C/s (10.3 °F/s). This value is the maximum safe cooling rate for this steel and the actual cooling rate cannot exceed this value. Preheating Temperature Requirement. Although the critical cooling rate cannot be exceeded, in the actual welding

operation a preheat can be used to reduce the cooling rate to 5.7 °C/s (10.3 °F/s). Assume that the welding condition is:

CURRENT (I), A ARC VOLTAGE (E), V HEAT-TRANSFER EFFICIENCY (η) TRAVEL VELOCITY (V), MM/S (IN./S) PLATE THICKNESS (T), MM (IN.)

250 25 0.9 7 (0.3) 9 (0.4)

The energy heat input, Hnet, is:

H net =

25(250)0.9 = 804 J / mm 7

(EQ 66)

Assuming that the thin-plate equation (Eq 50) applies: 2

3  R   9    = 32.2 = 0.0044   ( 550 − T0 )  2πλ  max  804 

(EQ 67)

resulting in a T0 of 162 °C (325 °F). The relative plate thickness should be checked:

τ =9

0.0044(550 − 162) = 0.41 804

(EQ 68)

Because τ is less than 0.75, the thin-plate equation does apply. If the initial plate temperature is raised either to or above 162 °C (325 °F), then the cooling rate will not exceed 5.7 °C/s (10.3 °F/s). Effect of Joint Thickness. If the plate thickness increases from 9 to 25 mm (0.36 to 1 in.), but there is the same level of

energy input, then the calculation of the initial plate temperature would be as follows. First, using the thin-plate equation (Eq 50): 2

3  R   25    = 32.2 = 0.0044   ( 550 − T0 )  2πλ  max  804 

(EQ 69)

resulting in a value for T0 of 354 °C (670 °F). The relative plate thickness, τ , should be checked:

τ = 25

0.0044(550 − 354) = 0.82 804

(EQ 70)

Because τ is greater than 0.75, the use of the thin-plate equation is inadequate. Using the thick-plate equation (Eq 49):

( 550 − T0 ) 32.2 =

2

(EQ 71)

804

resulting in a value for T0 of 389 °C (730 °F). The relative plate thickness should be checked:

τ = 25

0.0044(550 − 389) = 0.74 804

(EQ 72)

Although τ is less than, but near to, 0.75, using the thin-plate equation is adequate. Therefore, the initial temperature should be raised to 389 °C (730 °F) to avoid exceeding the cooling rate of 5.7 °C/s (10.3 °F). Now, if the plate thickness increases to 50 mm (2 in.), but there is the same level of energy input, then the thick-plate equation (Eq 49) applies and, again, the value for T0 is 389 °C (730 °F). The relative plate thickness should be checked:

τ = 50

0.0044(550 − 389) = 1, 48 804

Because τ is greater than 0.75, the use of the thick-plate equation is adequate.

(EQ 73)

Under some welding conditions, it is not necessary to reduce the cooling rate by using a preheat. For example, if the plate thickness is 5 mm (0.2 in.) and there is the same level of energy input: 2

3  R   5    = 32.2 = 0.0044   ( 550 − T0 )  2πλ  max  804 

(EQ 74)

resulting in a value for T0 of -24 °C (-11 °F). Therefore, using a preheat is unnecessary. Fillet-Welded "T" Joints. For a weld with a higher number of paths, as occurs in fillet-welded "T" joints, it is sometime

necessary to modify the cooling-rate equation, because the cooling of a weld depends on the available paths for conducting heat into the surrounding cold base metal. When joining 9 mm (0.35 in.) thick plate, where Hnet = 804 J/mm (20.4 kJ/in.), and when there are three legs instead of two, the cooling-rate equation is modified by reducing the effective energy input by a factor of

HNET =

2 3

(804) = 536 J/MM

2 : 3

(EQ 75)

Using the thin-plate equation (Eq 50): 2

3  R   9    = 32.2 = 0.0044   ( 550 − T0 )  2πλ  max  536 

(EQ 76)

resulting in a value for T0 of 254 °C (490 °F). The relative plate thickness should be checked:

τ =9

0.0044(550 − 254) = 0, 44 536

(EQ 77)

Because τ is less than 0.75, using the thin-plate equation is adequate. Therefore, a higher preheat temperature is more necessary than a butt weld because of the enhanced cooling. Example 3: Cooling Rate for the Location at Distance y (in cm) from the Centerline. For a steel plate of 25 mm (1 in.) thickness, (t), the welding condition is assumed to be:

HEAT INPUT (ηEI), KW (CAL/S) TRAVEL SPEED (V), CM/S (IN./S) PREHEAT (T0), °C (°F) NET ENERGY INPUT (HNET), CAL/CM

7.5 (1800) 0.1 (0.04) 20 (68) 18,000

The thermal properties needed for heat flow analysis are assumed to be:

MELTING TEMPERATURE (TM), °C (°F) THERMAL CONDUCTIVITY (λ), W/M · K (CALIT/CM · S · °C) SPECIFIC HEAT (CP), J/KG · C (BTU/LB · °F) DENSITY (ρ), G/CM3 (LB/IN.3)

1400 (2550) 43.1 (0.103) 473 (0.113) 7.8 (0.28)

Assume that one is interested in the critical cooling rate at the location on the surface (z = 0) at distance y = 2 cm from the centerline at the instant when the metal passes through the specific temperature of 615 °C (1140 °F). Initially, the relative plate thickness should be checked. From Eq 51, the relative plate thickness (using English units) is:

τ = 25

7.8 x0.113(615 − 20 = 4, 267 18, 000

Because τ is greater than 0.75, this plate can be treated as a thick plate. From Eq 78, cooling rate for thick plate at the location where the variables are w and r and at critical temperature θ= θc is:

∂θ ∂θ −vη EI  −v ( w + r )   − w v  w   = −v = exp   x  r ² − 2κ 1 + r   ∂t ∂w 2πλ r  2κ  

(EQ 78)

To solve Eq 78, we need to calculate the value of w and r first. From Eq 21, the temperature distribution of thick plate, and Eq 36 where z = 0, we can get:

θ − θ0 =

η EI  −v ( w + r )  exp   2πλ r 2κ  

(EQ 79)

and r= w²+y² By substituting the welding condition and material properties into Eq 21, Eq 36, and r = following simultaneous equations:

615 − 20 =

w² + y ² , we can obtain the

1800  −0.1( w + r )  exp  22π x0.103r  2 x0.117 

and r= w²+4 The value of w and r can be solved by using iteration techniques to solve the above simultaneous equation. The result is that w = -3 cm and r = 3.606 cm. Substituting w and r into Eq 78:

∂θ ∂θ −0.1x1800 0.1  −3    −0.1(−3 + 3.606   −(−3) = −v = x exp  x − 1 +   ∂t θ C ∂w 2π x0.103x3.606 2 x0.117    3.606² 2 x0.117  3.606   Therefore:

∂θ ∂t

= −9.78°C / s θC

From Eq 49 we can calculate the cooling rate along the centerline at the same temperature (615 °C, or 1140 °F):

R=

2π x0.103(615 − 20)² = 12.7°C / s 18, 000

Also from Eq 32, we can calculate the cooling rate in the heat-affected zone at a temperature of 615 °C (1140 °F):

  615 − 20 1.7  2  ∂θ −1  0.9843 − 0.486    x 1 tan = 0.35   +       0.125 ∂t θ C =615      π   336 / 2.362 

0.8

= 4°C / s

Therefore, at the same temperature, the cooling rate at the centerline is greater than the cooling rate at the location a distance y from the centerline. In addition, the cooling rate of the heat-affected zone is less than the cooling rate in the weld pool at the same temperature. Example 4: Solidification Rate. A weld pass of 800 J/mm (20.3 kJ/in.) in net energy input is deposited on a steel plate. The initial temperature is 25 °C (75 °F). The solidification time would be:

St =

2(800) = 0.94( s ) 2π (0.028)(0.0044)(1510 − 25)²

(EQ 80)

References cited in this section

22. HEAT FLOW IN WELDING, CHAPTER 3, WELDING HANDBOOK, VOL 1, 7TH ED., AWS, 1976 23. C.M. ADAMS, JR. COOLING RATE AND PEAK TEMPERATURE IN FUSION WELDING, WELD. J., VOL 37 (NO. 5), 1958, P 210S-215S 24. C.M. ADAMS, JR., COOLING RATES AND PEAK TEMPERATURES IN FUSION WELDING, WELD. J., VOL 37 (NO. 5), P 210-S TO 215-S 25. H. KIHARA, H. SUZUKI, AND H. TAMURA, RESEARCH ON WELDABLE HIGH-STRENGTH STEELS, 60TH ANNIVERSARY SERIES, VOL 1, SOCIETY OF NAVAL ARCHITECTS OF JAPAN, TOKYO, 1957 26. C.E. JACKSON, DEPARTMENT OF WELDING ENGINEERING, THE OHIO STATE UNIVERSITY LECTURE NOTE, 1977 27. C.E. JACKSON AND W.J. GOODWIN, EFFECTS OF VARIATIONS IN WELDING TECHNIQUE ON THE TRANSITION BEHAVIOR OF WELDED SPECIMENS--PART II, WELD. J., MAY 1948, P 253-S TO 266-S

Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

Parametric Effects

To show the effects of material property and welding condition on the temperature distribution of weldments, the welding of 304 stainless steel, low-carbon steel, and aluminum are simulated for three different welding speeds: 1.0, 5.0, and 8.0 mm/s (0.04, 0.02, and 0.03 in./s). The thermal conductivity and thermal diffusivity of 304 stainless steel are 26 W/m · K (0.062 cal/cm · s · °C) and 4.6 mm2/s (0.007 in.2/s), respectively. For low-carbon steel, the respective values are 50 W/m K (0.12 cal/cm s °C) and 7.5 mm2/s (0.012 in.2/s), whereas for aluminum, the respective values are 347 W/m K (0.93 cal/cm · s · °C) and 80 mm2 (0.12 in.2/s). The heat input per unit weld length was kept constant, 4.2 kJ/s (1 kcal/s), for all cases. The parametric results are described below. Effect of Material Type. Figures 5(a), 5(b), and 5(c) depict the effect of thermal properties on isotemperature contours

for a heat input of 4.2 kJ/s (1 kcal/s) and travel speeds of 1.0, 5.0, and 8.0 mm/s (0.04, 0.02, and 0.3 in./s). The temperature spreads over a larger area and causes a larger weld pool (larger weld bead) for low-conductivity material. The isotemperature contours also elongate more toward the back of the arc for low-conductivity material. For aluminum, a larger heat input would be required to obtain the same weld size as the stainless steel weldment.

FIG. 5(A) EFFECT OF THERMAL PROPERTY ON ISOTEMPERATURE CONTOURS FOR A HEAT INPUT OF 4.2 KJ/S (1000 CAL/S) AT A WELDING SPEED, V, OF 1 MM/S (0.04 IN./S) AND THE RESPECTIVE THERMAL CONDUCTIVITIES OF EACH MATERIAL (REFER TO TEXT FOR VALUES). VALUES FOR X AND Y ARE GIVEN IN CM, AND TEMPERATURES ARE GIVEN IN °C.

FIG. 5(B) EFFECT OF THERMAL PROPERTY ON ISOTEMPERATURE CONTOURS FOR A HEAT INPUT OF 4.2 KJ/S (1000 CAL/S) AT A WELDING SPEED, V, OF 5 MM/S (0.02 IN./S) AND THE RESPECTIVE THERMAL CONDUCTIVITIES OF EACH MATERIAL (REFER TO TEXT FOR VALUES). VALUES FOR X AND Y ARE GIVEN IN CM, AND TEMPERATURES ARE GIVEN IN °C.

FIG. 5(C) EFFECT OF THERMAL PROPERTY ON ISOTEMPERATURE CONTOURS FOR A HEAT INPUT OF 4.2 KJ/S (1000 CAL/S) AT A WELDING SPEED, V, OF 8 MM/S (0.3 IN./S) AND THE RESPECTIVE THERMAL CONDUCTIVITIES OF EACH MATERIAL (REFER TO TEXT FOR VALUES). VALUES FOR X AND Y ARE GIVEN IN CM, AND TEMPERATURES ARE GIVEN IN °C.

Welding Speed. Figures 6(a), 6(b), and 6(c) show the effect of welding speed on isotemperature contours. When the

travel speed increases, the weld size decreases and the isotemperature contours are more elongated toward the back of the arc. Larger heat inputs would be required for faster travel speeds in order to obtain the same weld size.

FIG. 6(A) EFFECT OF WELDING SPEED, V, ON ISOTEMPERATURE CONTOURS OF 304 STAINLESS STEEL FOR 4.2 KJ/S (1000 CAL/S) HEAT INPUT

FIG. 6(B) EFFECT OF WELDING SPEED, V, ON ISOTEMPERATURE CONTOURS OF LOW-CARBON STEEL FOR 4.2 KJ/S (1000 CAL/S) HEAT INPUT

FIG. 6(C) EFFECT OF WELDING SPEED, V, ON ISOTEMPERATURE CONTOURS OF ALUMINUM FOR 4.2 KJ/S (1000 CAL/S) HEAT INPUT

Figure 7 shows the effect of welding heat input on the peak temperature at two locations, 6.4 mm (

1 1 in.) and 13 mm ( 4 2

in.) from the weld centerline. (The material simulated in this illustration is low-carbon steel plate with a large thickness.) Within the practical range of welding conditions for the GMAW process of low-carbon steels, the peak temperature at both locations increases linearly, as the welding current increases, and decreases exponentially, with the travel speed. To •

increase welding current proportionally with travel speed for a constant heat input per unit weld length (that is. Q /V = constant), the peak temperature at two locations increases with the travel speed, which implies a larger weld size resulting from higher heat input. The welding current has a more significant effect on peak temperature than does the travel speed. However, the increase in peak temperature is more dominant at the 6.4 mm (

1 in.) location. The influence of 4

proportional increase in welding current and travel speed on peak temperature diminishes as the distance from the weld centerline increase and when the travel speed becomes high.

FIG. 7 PEAK TEMPERATURE DETERMINED BY THE POINT HEAT SOURCE SOLUTION. (A) AT 6.4 MM ( FROM THE CENTER. (B) AT 13 MM (

1 IN.) 2

1 IN.) FROM THE CENTER 2

Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

Thermophysical Properties of Selected Engineering Materials For the sake of convenience when using heat-flow equations, the thermal properties of selected engineering materials are provided in Table 2.

TABLE 2 THERMAL PROPERTIES OF SELECTED ENGINEERING ALLOYS ALLOY

CARBON STEEL ~0.5% C

DENSITY (ρ), AT 20 °C (68 °F), G/CM3

SPECIFIC HEAT (CP), AT 20 °C (68 °F), KJ/KG ·K

THERMAL CONDUCTIVITY (κ) AT 20 °C (68 °F, W/M · K)

COEFFICIENT OF THERMAL EXPANSION (α), AT 20 °C (68 °F), 10 -5/°C

7.833

0.465

54

1.474

~1.0% C ~1.5% C ALUMINUM AL-CU (94-96 % AL, 3-5% CU) AL-SI (86.5% AL. 1% CU) AL-SI (78-80% AL, 20-22% SI) AL-MG-SI (97% AL, 1% MG, 1% SI, 1% MN) TITANIUM ALLOY

STAINLESS STEEL CHROMIUM-NICKEL AUSTENITIC CHROMIUM FERRITIC CHROMIUM MARTENSITIC CARBON STEEL

7.801 7.753

0.473 0.486

43 36

1.172 0.970

2.787 2.659 2.627 2.707

0.883 0.867 0.854 0.892

164 137 161 177

6.676 5.933 7.172 7.311

4.500

0.52

16

0.84

DENSITY (ρ), AT 100 °C (212 °F), G/CM2

SPECIFIC HEAT (CP), AT 0-100 °C (32-212 °F), KJ/KG · K

THERMAL CONDUCTIVITY (κ), AT 100 °C (212 °F), W/M · K

COEFFICIENT OF THERMAL EXPANSION (α), AT 0538 °C (32-1000 °F), 10 -5/°C

7.800-8.000

0.46-0.50

18.7-22.8

1.700-1.920

7.800 7.800

0.46-0.50 0.42-0.46

24.4-26.3 28.7

1.120-1.210 1.160-1.210

7.800

0.48

60

1.17

Heat Flow in Fusion Welding Chon L. Tsai and Chin M. Tso, The Ohio State University

References

1. N.S. BOULTON AND H.E. LANCE-MARTIN, RESIDUAL STRESSES IN ARC WELDED PLATES, PROC. INST. MECH. ENG., VOL 33, 1986, P 295 2. D. ROSENTHAL AND R. SCHMERBER, THERMAL STUDY OF ARC WELDING, WELD. J., VOL 17 (NO. 4), 1983, P 2S 3. N.N. RYKALIN, "CALCULATIONS OF THERMAL PROCESSES IN WELDING," MASHGIZ, MOSCOW, 1951 4. K. MASUBUCHI, ANALYSIS OF WELDED STRUCTURES, PERGAMON PRESS, 1980 5. C.L. TSAI, "PARAMETRIC STUDY ON COOLING PHENOMENA IN UNDERWATER WELDING," PH.D THESIS, MIT, 1977 6. C.L. TSAI, FINITE SOURCE THEORY, MODELING OF CASTING AND WELDING PROCESSES II, ENGINEERING FOUNDATION MEETING, NEW ENGLAND COLLEGE (HENNIKER, NH), 31 JULY TO 5 AUG 1983, P 329 7. R. PAVELIC, R. TANAKUCHI, O. CZEHARA, AND P. MYERS, EXPERIMENTAL AND COMPUTED TEMPERATURE HISTORIES IN GAS TUNGSTEN ARC WELDING IN THIN PLATES, WELD. J., VOL 48 (NO. 7), 1969, P 295S 8. S. KOU, 3-DIMENSIONAL HEAT FLOW DURING FUSION WELDING, PROC. OF METALLURGICAL SOCIETY OF AIME, AUG 1980, P 129-138 9. P.G. KOGAN, THE TEMPERATURE FIELD IN THE WELD ZONE, AVE. SVARKA, VOL 4 (NO. 9), 1979, P 8 10. G.M. ECER, H.D. DOWNS, H.D. BRODY, AND M.A. GOKHALE, HEAT FLOW SIMULATION OF PULSED CURRENT GAS TUNGSTEN ARC WELDING, MODELING OF CASTING AND WELDING PROCESSES, ENGINEERING FOUNDATION 1980 MEETING (RINDGE, NH), 3-8 AUG 1980, P 139160

11. H. HIBBITT AND P. MARCAL, A NUMERICAL THERMOMECHANICAL MODEL FOR WELDING AND SUBSEQUENT LOADING OF A FABRICATED STRUCTURE, COMPUT. STRUCT., VOL 3, 1973, P 1145 12. E. FRIEDMAN, THERMOMECHANICAL ANALYSIS OF THE WELDING PROCESS USING FINITE ELEMENT METHODS, TRANS. ASME, AUG 1975, P 206 13. Z. PALEY AND P. HIBBERT, COMPUTATION OF TEMPERATURE IN ACTUAL WELD DESIGN, WELD. J., VOL 54 (NO. 11), 1975, P 385.S 14. T. NAKA, TEMPERATURE DISTRIBUTION DURING WELDING, J. JPN. WELD. SOC., VOL 11 (NO. 1), 1941, P 4 15. K. MASUBUCHI AND T. KUSUDA, TEMPERATURE DISTRIBUTION OF WELDED PLATES, J. JPN WELD. SOC., VOL 22 (NO. 5), 1953, P 14 16. C.L. TSAI AND C.A. HOU, THEORETICAL ANALYSIS OF WELD POLL BEHAVIOR IN THE PULSED CURRENT GTAW PROCESS, TRANSPORT PHENOMENA IN MATERIALS PROCESSING, ASME WINTER ANNUAL MEETING, 1983 17. E. FRIEDMAN, "FINITE ELEMENT ANALYSIS OF ARC WELDING," REPORT WAPD-TM-1438, DEPARTMENT OF ENERGY, 1980 18. J.S. FAN AND C.L. TSAI, "FINITE ELEMENT ANALYSIS OF WELDING THERMAL BEHAVIOR IN TRANSIENT CONDITIONS," 84-HT-80, ASME 19. R.L. APPS AND D.R. MILNER, HEAT FLOW IN ARGON-ARC WELDING, BR. WELD. J., VOL 2 (NO. 10), 1955, P 475 20. H.S. CARSLAW AND J.C. JAEGER, CONDUCTION OF HEAT IN SOLIDS, OXFORD PRESS 21. C.L. TSAI AND J.H. WU, "AN INVESTIGATION OF HEAT TRANSPORT PHENOMENA IN UNDERWATER WELDING," PRESENTED AT THE ASME WINTER ANNUAL MEETING (MIAMI BEACH, FL), 1985 22. HEAT FLOW IN WELDING, CHAPTER 3, WELDING HANDBOOK, VOL 1, 7TH ED., AWS, 1976 23. C.M. ADAMS, JR. COOLING RATE AND PEAK TEMPERATURE IN FUSION WELDING, WELD. J., VOL 37 (NO. 5), 1958, P 210S-215S 24. C.M. ADAMS, JR., COOLING RATES AND PEAK TEMPERATURES IN FUSION WELDING, WELD. J., VOL 37 (NO. 5), P 210-S TO 215-S 25. H. KIHARA, H. SUZUKI, AND H. TAMURA, RESEARCH ON WELDABLE HIGH-STRENGTH STEELS, 60TH ANNIVERSARY SERIES, VOL 1, SOCIETY OF NAVAL ARCHITECTS OF JAPAN, TOKYO, 1957 26. C.E. JACKSON, DEPARTMENT OF WELDING ENGINEERING, THE OHIO STATE UNIVERSITY LECTURE NOTE, 1977 27. C.E. JACKSON AND W.J. GOODWIN, EFFECTS OF VARIATIONS IN WELDING TECHNIQUE ON THE TRANSITION BEHAVIOR OF WELDED SPECIMENS--PART II, WELD. J., MAY 1948, P 253-S TO 266-S Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Introduction MOLTEN WELD POOLS are dynamic. Liquid in the weld pool in acted on by several strong forces, which can result in high-velocity fluid motion. Fluid flow velocities exceeding 1 m/s (3.3 ft/s) have been observed in gas tungsten arc (GTA) welds under ordinary welding conditions, and higher velocities have been measured in submerged arc welds. Fluid flow is important because it affects weld shape and is related to the formation of a variety of weld defects. Moving liquid transports heat and often dominates heat transport in the weld pool. Because heat transport by mass flow depends on the

direction and speed of fluid motion, weld pool shape can differ dramatically from that predicted by conductive heat flow. Temperature gradients are also altered by fluid flow, which can affect weld microstructure. A number of defects in GTA welds have been attributed to fluid flow or changes in fluid flow, including lack of penetration, top bead roughness, humped beads, finger penetration, and undercutting. Instabilities in the liquid film around the keyhole in electron beam and laser welds are responsible for the uneven penetration (spiking) characteristic of these types of welds. Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Mass Transport in the Arc High-velocity gas motion occurs in and around the arc during welding. The gas motion is partially due to cover gas flow, but, more importantly, it is driven by electromagnetic forces associated with the arc itself. In gas metal arc (GMA) welds, liquid filler metal is also being transferred through the arc from the electrode to the workpiece. Both the mode and velocity of metal transfer have major effects on weld pool shape. Mass transport in and around the arc is important in GTA welding (GTAW) and even more so in GMA welding (GMAW); a detailed description of this phenomenon can be found in the articles on arc physics and metal transfer and weld behavior in this Volume. Only the effects of this transport on the weld pool will be discussed in this article. Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Gas Tungsten Arc Welding Most experimental and theoretical work on weld pool fluid flow and its effects has been directed toward GTAW. The motivation for much of this work was the observation of dramatically different weld pool shapes for GTA welds made using identical welding parameters on different heats of the same material with the same nominal composition. An extreme example of weld shape variability is shown in Fig. 1. Early observations of variable weld shape (often referred to as variable penetration) were not only an intellectual puzzle but also an indication of a growing practical problem. Gas tungsten arc welding is commonly used for high-precision, high-quality automated welding applications, where reproducibility of weld shape or penetration is critical.

FIG. 1 PARTIAL-PENETRATION GTA WELDS MADE UNDER THE SAME WELDING CONDITIONS ON TWO HEATS OF TYPE 304L STAINLESS STEEL HAVING THE SAME NOMINAL COMPOSITION. (A) 3 PPM S, D/W = 0.2. (B) 160 PPM S, D/W = 0.40. 9×

The possibility that fluid flow in the weld pool could alter weld shape has been recognized for many years. For example, in 1965 Christensen et al. (Ref 1) proposed that convection is partially responsible for deviations in weld pool shapes from those predicted by conduction solutions. The forces driving fluid flow in GTA weld pools have also been long known. The four primary driving forces are surface tension gradients, electromagnetic or Lorentz forces, buoyancy forces, and aerodynamic drag forces caused by passage of the arc plasma over the weld pool surface.

Surface-Tension-Driven Fluid Flow Model. Surface tension gradients were first proposed by Ishizake et al. (Ref 2) as

potential driving forces for weld pool fluid flow. Surface-tension-driven fluid flow was first described by Thomson (Ref 3) in 1855, but the phenomenon is commonly called Marangoni convection from the work of Carlo Marangoni (Ref 4). In 1982, Heiple and Roper (Ref 5) proposed that surface tension gradients are commonly the dominant forces driving fluid flow in GTA welds and that these gradients could be drastically altered by very small concentrations of certain trace elements. Surface tension gradients exist on a weld pool surface because the surface tension is temperature dependent, and there are large temperature gradients on a weld pool surface. For pure metals and many alloys, the surface tension decreases as temperature increases; that is, the surface tension temperature coefficient is negative. For weld pools in such materials, the surface tension will be greatest on the coolest part of the pool surface at the edge and lowest on the hottest part under the arc near the center of the pool. Such a surface tension gradient produces outward surface fluid flow, as shown schematically in Fig. 2(a). This fluid flow pattern transfers heat efficiently from the hottest part of the weld pool (near the center) to the edge and produces a relatively wide and shallow weld.

FIG. 2 SCHEMATIC SHOWING SURFACE FLUID FLOW (TOP) AND SUBSURFACE FLUID FLOW (BOTTOM) IN THE WELD POOL. (A) NEGATIVE SURFACE TENSION TEMPERATURE COEFFICIENT (PURE MATERIAL). (B) POSITIVE SURFACE TENSION TEMPERATURE COEFFICIENT (SURFACE-ACTIVE ELEMENTS PRESENT)

Certain elements are surface active in molten metals; that is, they segregate to the surface of the solvent liquid metal and lower the magnitude of the surface tension, often drastically. Small concentrations of surface-active additions can also change the temperature dependence of the surface tension of the solvent metal or alloy so that, for a limited range of temperature above the melting point, the surface tension increases with increasing temperature. With a positive surface tension temperature coefficient, the surface tension will be highest near the center of the weld pool. Such a surface tension gradient will produce fluid flow inward along the surface of the weld pool and then down, as indicated schematically in Fig. 2(b). This fluid flow pattern transfers heat efficiently to the bottom of the weld pool and produces a relatively deep, narrow weld. Experimental Observations. This physical model was developed and verified with a series of experiments in which stainless steel base metal was doped with low concentrations of various elements and the effect of the doping on weld pool shape measured. High-speed motion pictures of the weld pool surface suggested the fluid flow patterns indicated in Fig. 2. The addition of sulfur, oxygen, selenium, and tellurium to stainless steel in low concentrations (less than 150 ppm) was shown to substantially increase GTA weld depth-to-width ratio (d/w). All these elements are known to be highly surface active in liquid iron. Measurements of the temperature dependence of the surface tension for steels with different GTA weld penetration characteristics produced an impressive correlation between a positive surface tension temperature coefficient arising from surface-active impurities and high d/w ratio welds (Fig. 3).

FIG. 3 PLOT OF SURFACE TENSION VERSUS TEMPERATURE FOR TWO LIQUID STEELS. THE DATA LABELED "HIGH D/W HEAT" ARE FROM MATERIAL HAVING APPROXIMATELY 160 PPM MORE SULFUR THAN THE MATERIAL LABELED "LOW D/W HEAT." THE DASHED LINES INDICATE THE EXPECTED BEHAVIOR OF THE SURFACE TENSION ABOVE THE MAXIMUM TEMPERATURE STUDIED. SOURCE: REF 6

When additions were made to the weld pool of elements known to react with surface-active elements already present in the steel to form compounds that are not surface active, the GTA weld d/w ratio decreased. Aluminum reacts with oxygen and produced wider, shallower welds in 21-6-9 stainless steel. Cerium reacts with both sulfur and oxygen and also produced lower d/w ratio welds. The effects of trace elements on weld shape have also been observed in a number of other alloys. These observations are summarized in Table 1.

TABLE 1 EFFECT OF TRACE ELEMENT IMPURITIES ON GTA WELD PENETRATION OF SELECTED ALLOYS

ALLOY SYSTEM

ZIRCALOY-2 IRON-BASE ALLOYS STAINLESS STEELS 304, 316, 21-6-9 JBK-75 AISI 8630 2.5CR1MO NICKEL-BASE ALLOYS INCONEL 600, 718 (A) (B)

TRACE ELEMENT IMPURITY INCREASES DECREASES WELD PENETRATION WELD PENETRATION CHLORINE ... SULFUR, OXYGEN, CALCIUM, ALUMINUM, CERIUM(A), SELENIUM(A), LANTHANUM(A), SILICON(B), TELLURIUM(A) TITANIUM(B)

SULFUR

OXYGEN

UNCOMMON IMPURITIES EFFECT NEGLIGIBLE OR UNCERTAIN

The simple physical model illustrated in Fig. 2 has been remarkably successful in qualitatively explaining trace element effects. For example, changing GTA welding conditions alters the magnitude and distribution of arc energy input to the weld, which in turn changes temperature gradients on the weld pool surface. From Fig. 2, a change in welding conditions that makes the center of the weld hotter, such as increasing current, should drive the existing fluid flow pattern more strongly. As shown in Fig. 4, increasing current improves the d/w ratio of steel doped with surface-active elements and reduces it for high-purity base metal. If the center of the weld becomes so hot that there is a region where the temperature coefficient of the surface tension is no longer positive, then the fluid flow pattern necessary for deep penetration is disrupted and the d/w ratio decreases. This effect is seen at high currents in Fig. 4. Similar results have been obtained for other welding parameters. The surface temperature at which the change from positive to negative surface tension

temperature coefficient occurs for stainless steel is estimated by extrapolation in Fig. 3 to be about 2050 °C (3720 °F). Detailed thermodynamic calculations of the temperature dependence of the surface tension of iron-sulfur alloys predict that the transition from positive to negative dγ/dT will occur at 2032 °C (3690 °F) for the high-sulfur alloy in Fig. 3. Recent spectrographic weld pool temperature measurements and numerical simulations have indicated that this temperature can be exceeded in stainless steel GTA weld pools under normal welding conditions.

FIG. 4 PLOT OF WELD D/W RATIO VERSUS WELD CURRENT FOR THE STARTING BASE METAL (TYPE 304 STAINLESS STEEL WITH VERY LOW RESIDUAL IMPURITY CONTENT) AS WELL AS FOR ZONES DOPED WITH SULFUR AND SELENIUM. SOURCE: REF 7

The surface-tension-driven fluid flow model should be applicable to non-arc processes, provided the energy input distribution is similar to a GTA arc. This condition is satisfied for conduction-mode electron beam and laser welds. Dramatic increases in weld d/w ratio in selenium-doped zones in stainless steel have been observed for both traveling laser and electron beam conduction-mode welds. The weld shape changes were similar to those observed for GTA welds. Conduction-mode electron beam welds can also be used to demonstrate that variations in weld shape with changes in welding parameters, as illustrated in Fig. 4, are not a result of some complex arc/weld pool interaction. One of the results of an investigation of the effect of changes in beam focus on weld shape in electron beam welds on low- and high-sulfur materials is shown in Fig. 5. The high-sulfur material exhibits a maximum in d/w ratio with increasing power density at a moderate power density away from sharp focus, which is analogous to that shown with increasing current in Fig. 4. (The low- and high-sulfur stainless steels have deep and nearly identical penetration near sharp focus. Near sharp focus, where the power density is high, penetration is by a keyhole mechanism in which average penetration is unrelated to surfacetension-driven fluid flow.) Measurements of the electron beam power-density distribution verified that there were no anomalous changes in the beam, such as a beam width maximum, with increasing peak power density. The d/w ratio maximum away from sharp focus is therefore proposed to originate from exactly the same mechanism as for increasing current with GTA welds.

FIG. 5 PLOT OF ELECTRON BEAM WELD POOL RATIO (D/W) VERSUS ELECTRON BEAM POWER DENSITY FOR LOW-SULFUR (20 PPM) AND HIGH-SULFUR (>120 PPM) TYPE 304L STAINLESS STEEL. KEYHOLE FORMATION BEGINS AT ABOUT 2 × 103 W/MM2.

Measurements of electron beam power distribution were made as a function of beam focus and were used to calculate the beam power density at the d/w maximum. The power density was also calculated at the d/w maximum for GTA welds. Calculation of weld pool surface temperatures using a traveling distributed heat source model showed the central surface temperatures to be essentially identical at the d/w maxima for the electron beam and GTA processes. The calculated surface temperatures using the traveling distributed heat source conduction approximation are much too high because, as indicated previously, most of the heat transport in the weld pool is by mass flow rather than by conduction. Nevertheless, the equality of the calculated weld pool peak temperatures at the d/w maxima provides strong confirmation that the mechanism responsible for the presence of the maximum in weld d/w with increasing input power density is independent of heat source and is not a result of an arc phenomenon. A final example of the success of the fluid flow model in explaining GTA welding phenomena is provided by butt welding together two steels with large differences in weld penetration characteristics. The weld pool is not centered over the joint; rather, it is displaced toward the material with low d/w behavior (Ref 8), as indicated schematically in Fig. 6. The low d/w material has a low concentration of surface-active impurities and therefore a high surface tension. Thus, there is a net surface tension gradient across the weld pool toward the low d/w material, producing the fluid flow pattern and weld cross section indicated in Fig. 6. The actual fluid flow pattern is certain to be more complicated than that illustrated.

FIG. 6 SCHEMATIC SHOWING TYPICAL FLUID FLOW GENERATED WHEN BUTT WELDING TWO HEATS OF MATERIAL WITH DIFFERENT PENETRATION CHARACTERISTICS. SOURCE: REF 8

Numerical Simulations. Direct measurement of weld pool fluid flow is very difficult. Surface flow has been studied by

observing the motion on the surface of slag or intentionally added particles, as has bulk flow in transparent liquid/solid systems using simulated welding heat sources. There are a number of numerical calculations of weld pool fluid flow and shape, the most comprehensive of which appears to be that of Zacharia et al. (Ref 9). These numerical fluid flow simulations agree in broad terms with the physical model illustrated in Fig. 2; however, they differ in detail with Fig. 2 and with one another. One result of the analysis of Zacharia is that the weld pool surface temperature reached in stationary welds exceeds that in traveling welds. Trace element effects should therefore be less pronounced in stationary welds, because a larger portion of the pool surface is above the temperature range where a positive surface tension temperature coefficient exists. This prediction is in agreement with experimental observations. Interactions. Experiments in which low concentrations of various elements have been added to the base metal have

demonstrated clear changes in weld pool shape. However, when weld d/w ratio is plotted versus chemical analysis for numerous heats of material, considerable scatter is observed. A comparison of weld d/w ratio in a standard weldability test versus sulfur content for about 200 lots of steel is shown in Fig. 7. There is a clear trend of increasing d/w with increasing sulfur, but the variability for a given sulfur content is substantial. Some of this variability is associated with imprecision in chemical analysis for sulfur and some with variations in oxygen content. However, it appears that much of the variability is related to interactions of the surface-active elements sulfur and oxygen with other components of the steel. Calcium is known to react with oxygen and, to a lesser extent, with sulfur to form stable compounds unlikely to be surface active. Aluminum and silicon also react with oxygen to form stable compounds. Thus, the amount of sulfur and oxygen available for segregation to the weld pool surface is a complicated function of the total weld pool chemistry.

FIG. 7 PLOT OF WELD D/W RATIO VERSUS SULFUR CONTENT FOR APPROXIMATELY 200 HEATS OF TYPE 304L STAINLESS STEEL. EACH POINT IS AN AVERAGE OF MULTIPLE SULFUR ANALYSES AND WELD D/W RATIO MEASUREMENTS. IF SINGLE VALUES ARE USED, THE SCATTER IS GREATER. SOURCE: REF 10

There is an additional complication, illustrated in Fig. 8. When oxygen is added to the shielding gas, weld d/w ratio increases, passes through a maximum, and then declines with increasing oxygen content. Similar effects are seen with SO2 additions. The surface of welds made with torch gas concentrations above the d/w maximum are heavily oxidized. A possible explanation for the decreasing d/w ratios at higher oxygen concentrations is that a liquid oxide film (slag) is formed on the weld pool surface, altering the surface tension gradients.

FIG. 8 PLOT OF WELD D/W RATIO VERSUS OXYGEN CONCENTRATION IN THE TORCH GAS FOR GTA BEAD-ONPLATE WELDS ON 21-6-9 STAINLESS STEEL. SOURCE: REF 11

Oxygen can be added to the weld pool in other ways. Oxidizing the plate surface prior to welding adds oxygen to the weld pool and increases the weld d/w ratio. Wire brushing and grit blasting an originally clean surface increases both the surface oxide thickness and the surface area, thereby adding oxygen to the weld pool. The d/w ratio of JBK-75 stainless steel (a modification of ASTM A-286) GTA welds as a function of amount of wire brushing is similar in form to Fig. 8. Welds on plate brushed beyond the weld d/w maximum had extensive slag on the weld surface. The weld d/w ratio on JBK-75 stainless steel appears to be particularly sensitive to oxygen additions. Similar wire-brushing experiments on type 304 stainless steel showed smaller effects. Although the origin of the difference in sensitivity between JBK-75 and type 304 is not known, one possibility is indicated by surface tension measurements on iron-silicon alloys in contact with carbon dioxide. The effect of CO2 on the surface tension was a strong function of silicon content. For low-silicon alloys, the surface tension dropped sharply when contacted with CO2. For alloys with more than 1.2% Si, the surface tension increased when contacted with CO2. The different behavior was attributed to differences in slag formation on the liquid metal surface. Thus, silicon appears to interfere with the ability of oxygen to produce a positive surface tension temperature coefficient on liquid iron. The JBK75 stainless steel used in the wire-brushing experiments contained only 0.06% Si, and the 21-6-9 used for the torch gas experiment (Fig. 8) contained only 0.16% Si. Type 304 stainless steel typically contains more than 0.5% Si. Thus, the high sensitivity of JBK-75 to oxygen additions may be related to its unusually low silicon content. Reports in the literature on the effect of oxygen on GTA weld penetration have been somewhat variable; differences in weld pool silicon content may be partially responsible for this variability. Effects of High Current. Electromagnetic (Lorentz) stirring of the weld pool becomes more important at higher

currents. The direction of the Lorentz force produces the deep penetration fluid flow pattern indicated in Fig. 2(b). At sufficiently high current, the Lorentz force dominates other forces that drive fluid flow, and the effects of trace elements become less important. In addition, the plasma jet becomes stronger with increasing current and at high enough currents produces a significant depression of the weld pool surface. The plasma jet forces are resisted by surface tension forces on the weld pool. Halmoy (Ref 12) has shown that for a traveling weld the radial pressure gradient from the plasma jet tends to transport liquid from the front to the rear of the weld (Fig. 9). For a sufficiently strong pressure gradient, pB, the liquid level under the arc may be pushed down to the bottom of the pool, as shown in Fig. 9(c). High-speed motion pictures of the effect of sulfur additions on weld pool fluid flow in 21-6-9 stainless steel showed behavior similar to that in Fig. 9(c) after large sulfur additions. Sulfur additions substantially reduce the weld pool surface tension and thereby increase the effect of the plasma jet.

FIG. 9 EFFECT OF ARC PRESSURE ON THE WELD POOL FOR STATIONARY AND TRAVELING WELDS. (A) V = 0. (B) V > 0, WITH WEAK RADIAL PRESSURE GRADIENT, P1. (C) V > 0, WITH STRONG RADIAL PRESSURE GRADIENT, PB. SOURCE: REF 12

In addition, a surface depression almost certainly changes the energy distribution input to the weld pool by the arc compared with a flat weld pool surface. Finally, under some high-current conditions, a vortex has been observed near the center of weld pools. Strategies for Controlling Poor and Variable Penetration. A number of techniques have been developed to reduce

penetration variability and improve penetration. Sometimes the weld pool shift illustrated in Fig. 6, which occurs when heats with different penetration characteristics are welded together, can be minimized by very tight heat sinking. The sensitivity of weld shape to trace element differences is a function of welding parameters, as illustrated in Fig. 4 for current; thus, combinations of welding parameters can be chosen that minimize heat-to-heat penetration variability. For wire-fed joints, some joint designs are more tolerant of penetration variability than others. For example, joints with thinner, wider lands are less sensitive than narrower grooves with thicker lands. Material control by selection or specification is another approach. It was originally anticipated that stainless steels with less than 20 ppm S would have poor, but consistent, penetration. However, the data illustrated in Fig. 7 demonstrate that penetration variability is substantial even at these very low sulfur levels. Stainless steels with more than about 100 ppm S generally have consistently good penetration. A related approach is to test incoming material for welding characteristics and then select material with desired welding behavior for critical applications. Finally, the weld pool can be doped with enough surface-active elements to ensure good penetration. Sulfur and oxygen are the most practical dopants for ferrous alloys. The addition of oxygen or sulfur dioxide to the torch shielding gas has been shown to improve penetration and reduce penetration variability; however, this approach presents several practical problems. Oxygen can also be added to the weld pool by oxidizing the weld joint or by otherwise increasing the oxygen content of the weld groove surface--by wire brushing, for example. Probably the most useful approach is to dope the weld pool in the wire-fed joints by using a special filler wire. A group of defense contractors had a special heat of type 380L stainless steel produced with a sulfur content in the range of 100 to 150 ppm. The heat was then converted into weld wire. Welds made with this wire on a variety of stainless steel base metals have exhibited consistent and good penetration.

References cited in this section

1. N. CHRISTENSEN, V. DE L. DAVIES, AND K. GJERMUNDSEN, DISTRIBUTION OF TEMPERATURES IN ARC WELDING BR. WELD.J., VOL 12, 1965, P 54-75 2. K. ISHIZAKE, K. MURAI, AND Y. KANBE, "PENETRATION IN ARC WELDING AND CONVECTION IN MOLTEN METAL," DOCUMENT 77-66, INTERNATIONAL INSTITUTE OF WELDING, STUDY GROUP 212, 1966 3. J. THOMSON, ON CERTAIN CURIOUS MOTIONS OBSERVABLE AT THE SURFACES OF WINE AND OTHER ALCOHOLIC LIQUORS, PHILOS.MAG., VOL 10, 1855, P 330-333 4. C. MARANGONI, ÜBER DIE AUSBREITUNG DER TROPFEN EINER FLUESSIGKEIT AUF DER OBERFLAESCHE EINER ANDEREN, ANN. PHYS. CHEM., VOL 143 (NO. 7), 1871, P 337-354

5. C.R. HEIPLE AND J.R. ROPER, MECHANISM FOR MINOR ELEMENT EFFECT ON GTA FUSION ZONE GEOMETRY, WELD. J., VOL 61, 1982, P 97S-102S 6. B.J. KEENE, K.C. MILLS, AND R.F. BROOKS, SURFACE PROPERTIES OF LIQUID METALS AND THEIR EFFECTS ON WELDABILITY, MATER. SCI.TECHNOL., VOL 1, 1985, P 568-571 7. P. BURGARDT AND C.R. HEIPLE, INTERACTION BETWEEN IMPURITIES AND WELDING PARAMETERS IN DETERMINING GTA WELD SHAPE, WELD. J., VOL 65, 1986, P 150S-155S 8. M.J. TINKLER, I. GRANT, G. MIZUNO, AND C. GLUCK, WELDING 304L STAINLESS STEEL TUBING HAVING VARIABLE PENETRATION CHARACTERISTICS, THE EFFECTS OF RESIDUAL, IMPURITY, AND MICRO-ALLOYING ELEMENTS ON WELDABILITY AND WELD PROPERTIES, P.H.M. HART, ED., PAPER 29, THE WELDING INSTITUTE, CAMBRIDGE, 1984 9. T. ZACHARIA, S.A. DAVID, J.M. VITEK, AND T. DEBROY, WELD POOL DEVELOPMENT DURING GTA AND LASER BEAM WELDING OF TYPE 304 STAINLESS STEEL, PART 1-THEORETICAL ANALYSIS, WELD. J., VOL 68, 1989, P 499S-509S 10. P. BURGARDT AND R.D. CAMPBELL, CHEMISTRY EFFECTS ON STAINLESS STEEL WELD PENETRATION, FERROUS ALLOY WELDMENTS, TRANS TECH PUBLICATIONS, SWITZERLAND, 1992, P 379-415 11. C.R. HEIPLE, P. BURGARDT, AND J.R. ROPER, THE EFFECT OF TRACE ELEMENTS ON GTA WELD PENETRATION, MODELING OF CASTING AND WELDING PROCESSES II, J.A. DANTZIG AND J.T. BERRY, ED., TMS-AIME, 1984, P 193-205 12. E. HALMOY, THE PRESSURE OF THE ARC ACTING ON THE WELD POOL, ARC PHYSICS AND WELD POOL BEHAVIOR, THE WELDING INSTITUTE, CAMBRIDGE, 1979, P 259-266

Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Deep-Penetration Electron Beam and Laser Welds Keyhole Formation. A fundamental difference between arc heat sources and electron or laser beam heat sources is that electron and laser beams are capable of delivering heat over a small area at much higher power densities. As the power density of the welding heat source is increased, the peak surface temperature of the weld pool rises. For many metals, vapor pressure rises nearly exponentially with temperature and becomes appreciable (above 4000 Pa, or 30 torr, for an electron beam operating in a vacuum) near 0.8 Tb, where Tb is the material boiling point in degrees Kelvin. As the surface temperature approaches this value, the liquid surface under the power source is depressed by the vapor pressure. As the liquid moves away from the power source, the surface is depressed and a cavity is formed. This is the basic theory of keyhole formation by an electron or laser beam. Fluid Flow in the Keyhole. The keyhole is a cavity having roughly the size and shape of the beam; that is, it is usually

approximately cylindrical. Fluid flow in the thin layer around the cavity plays a major role in determining the behavior of deep-penetration, keyhole-mode electron beam and laser welds. The first flow required for a traveling weld is transport of metal melted at the front wall of the cavity to the rear (where it eventually solidifies). Because of the weld geometry, this liquid must move around the keyhole cavity as a thin, high-velocity layer on the walls of the cavity. The dominant driving force for this motion appears to be surface tension gradients. As in GTAW, these gradients arise because the surface tension is temperature dependent and there is a substantial temperature difference between the front and rear of the cavity. Calculations of the fluid flow have been performed by Wei and Giedt (Ref 13). Their calculations for pure iron predict that liquid metal is moved from the front to the rear of a sharp focus keyhole in a film that is about 0.02 mm (0.0008 in.) thick and is moving at a velocity of about 250 mm/s (10 in./s). The fluid is driven by a temperature difference of about 400 °C (720 °F). These results are only nominal values from the calculations, but they change little with the assumptions used and are representative of electron beam welds in general.

There is an interesting potential problem with this model. As discussed previously, the surface tension temperature coefficient, dγ/dT, can be positive over a limited temperature range in steels and some other alloys if surface-active impurities are present in sufficient quantity. If dγ/dT is positive, then it would appear that molten metal would not be transported around the cavity. The likely resolution of this difficulty is that the liquid dwells slightly longer in the front of the cavity and is heated by the beam above the temperature at which a transition to negative dγ/dT occurs. Some perturbation of the normal fluid flow may occur under these circumstances. Effects of such a perturbation on weld characteristics appear to be uncommon. The only available evidence for such an effect comes from electron beam welds performed on a heat of type 304 stainless steel with very low residual impurity content, except for about 340 ppm S. These welds had very high porosity compared with identical welds in other heats of material. Instability in Keyhole Fluid Flow. In addition to the steady-state flow of liquid around the keyhole, significant

instability in the fluid motion has been seen. For example, Mara et al. (Ref 14) used side-view, self-illuminated X-ray films to show that the beam locations shifts between full penetration and nearly zero penetration in an irregular fashion. High-speed X-ray photographs show that the variable penetration is caused by a lump of metal that sags into the keyhole from high up on the cavity rear wall. The sequence of events that creates weld penetration irregularity was analyzed by Tong and Giedt (Ref 15) and is illustrated schematically in Fig. 10. First, the beam forms a keyhole and produces a lump of very hot displaced metal at the top rear of the traveling cavity (Fig. 10a). After some period of time, the weld reaches the full penetration allowed by heat flow (Fig. 10b). The lump of displaced metal is unstable and eventually falls into the cavity, partially filling it. The beam must now drill through this additional material. The sequence (a) to (c) in Fig. 10 occurs repeatedly, but irregularly, and produces weld penetration variability. When particularly severe, the irregular penetration is called spiking. Spiking also leads to a tendency for voids to become trapped in the root of the weld.

FIG. 10 SCHEMATIC SHOWING KEYHOLE INSTABILITY. (A) KEYHOLE IS FORMED BY HEAT GENERATED BY ELECTRON BEAM. (B) MAXIMUM PENETRATION THAT CAN BE PRODUCED BY HEAT FLOW. (C) LIQUID COOLS, CAUSING IMPENDING COLLAPSE OF DISPLACED METAL. THE KEYHOLE IS FILLED BY A LUMP OF COOLING MATERIAL AT THE END OF (C), WHICH RETURNS THE KEYHOLE TO CONDITION (A) TO RESTART THE SEQUENCE.

A mathematical model that is the basis for understanding this keyhole instability was developed by Giedt et al. (Ref 13, 15). Stability of the keyhole is basically a balance between vapor pressure, which keeps the cavity open, and surface tension, which tries to close the cavity. Vapor pressure and surface tension are both functions of temperature. Vapor pressure increases nearly exponentially with temperature for many materials, while surface tension generally decreases slowly with temperature. Because there is a substantial temperature gradient from the top to the bottom of the keyhole, both vapor pressure and surface tension vary with depth in the keyhole. Direct measurements of cavity wall temperatures for electron beam welds indicate that the temperature at the top is nearly equal to the melting temperature, while the temperature at the bottom is high enough that the vapor pressure there is large. In type 304 stainless steel, the temperature at the bottom of the cavity is about 2200 °C (1200 °F). The data indicate that the temperature profile is insensitive to welding variables and is determined primarily by material properties, that is, by the temperature where the vapor pressure exceeds about 1330 Pa (10 torr).

Based on temperature measurements of a typical keyhole, the vapor pressure decreases approximately exponentially with height from the bottom of the keyhole and approaches zero at the top of the keyhole. The vapor pressure is balanced primarily by surface tension. For pure iron and many other materials, the surface tension decreases approximately linearly with temperature. Thus, the inward surface tension pressure increases slowly from the bottom to the top of the keyhole. The balance of vapor pressure and surface tension is such that the cavity is stable near the bottom, where the vapor pressure tending to expand the cavity exceeds the surface tension tending to collapse it. At the top, the cavity is unstable because the surface tension tending to collapse it exceeds the vapor pressure tending to expand it. Thus, the cavity will always tend to be filled by liquid originating from above some height H in the keyhole, the pressure crossover height. This cavity filling produces weld penetration irregularity. Weld penetration irregularity is a consequence of the nature of the keyhole in deep-penetration welding and will not generally be solved by simple equipment or material modifications. Penetration irregularity is usually less severe if H (as measured from the top) is a small fraction of the keyhole depth. In such cases the volume of the lump is small, and liquid tends to flow into the cavity relatively smoothly. If H is a large fraction of the keyhole depth, then the liquid lump has a relatively large volume and, when it falls into the keyhole at irregular times, it nearly fills the cavity and produces large penetration variations.

References cited in this section

13. P.S. WEI AND W.H. GIEDT, SURFACE TENSION GRADIENT-DRIVEN FLOW AROUND AN ELECTRON BEAM WELDING CAVITY, WELD. J.,VOL 64, 1985, P 251S-259S 14. G.L. MARA, E.R. FUNK, R.C. MCMASTER, AND P.E. PENCE, PENETRATION MECHANISMS OF ELECTRON BEAM WELDING AND THE SPIKING PHENOMENON, WELD. J., VOL 53, 1974, P 246S-251S 15. H. TONG AND W.H. GIEDT, A DYNAMIC INTERPRETATION OF ELECTRON BEAM WELDING, WELD. J., VOL 49, 1970, P 259S-266S Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Gas Metal Arc Welding Fluid flow certainly occurs in GMA weld pools, but reports on the details of its nature and effects are quite limited. In the spray transfer mode, the impact of the stream of droplets from the electrode on the weld pool forms a substantial depression or crater. This is the mechanism responsible for the typical fingerlike penetration observed with argon as the torch gas. Thus, the depth of penetration is primarily dependent on the momentum of the stream of droplets. A plot of penetration versus the momentum of the droplet stream shows an excellent correlation (Fig. 11). Because the major force driving the droplets toward the weld is drag from the plasma jet, the kinematic viscosity (density times viscosity) of the shielding gas is important. For example, argon has a substantially higher kinematic viscosity than helium, so argon drives the droplet stream more effectively, resulting in a deeper, narrower penetration finger than with helium torch gas. The strength of the plasma jet can also be altered by changing the ambient pressure. For GMA welds on aluminum. Amson and Salter (Ref 17) saw a steady decrease in penetration with pressure until at 13,300 Pa (100 torr) there was essentially no penetration at all. Another way to modify the penetration depth is to spread the stream of droplets over the surface of the weld. This can be accomplished by applying a varying transverse magnetic field.

FIG. 11 PLOT OF GMA WELD PENETRATION IN MILD STEEL FOR SPRAY TRANSFER MODE VERSUS MOMENTUM OF DROPLET STREAM. SOURCE: REF 16

References cited in this section

16. W.G. ESSERS AND R. WALTER, HEAT TRANSFER AND PENETRATION MECHANISMS WITH GMA AND PLASMA-GMA WELDING, WELD. J., VOL 62, 1981, P 37S-42S 17. J.C. AMSON AND G.R. SALTER, ANALYSIS OF THE GAS-SHIELDED CONSUMABLE METALARC WELDING SYSTEM, BR. WELD. J., VOL 10, 1963, P 472-483 Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Submerged Arc Welding Currents commonly employed in submerged arc welding (SAW) are much higher than those used in GTAW or GMAW. Submerged arc welding currents often exceed 1000 A. Thus, the electromagnetic or Lorentz force combined with the tendency of the radial pressure gradient in the moving cavity to transport liquid to the rear of the cavity (Fig. 9) are likely the dominant forces driving fluid flow. The generally accepted flow pattern is indicated schematically in Fig. 12. A cavity is formed at the front of the moving weld pool by arc pressure and by the momentum of drops from the rapidly melting electrode. Metal that is melted at the front of the pool flows underneath and on either side of the cavity. At the rear of the pool, the flow reverses and metal flows back toward the cavity along and near the surface. Flow velocities can be very high; Eichhorn and Engel (Ref 19) measured 4 m/s (13 ft/s) at 720 A. The flow pattern was derived from observations of motion of marker elements, added to the weld pool, as determined from subsequent metallographic sections, and from the motion of radioactive tracer additions. The general features illustrated in Fig. 12 appear to be supported by X-ray fluoroscopy observations. Major perturbations in the indicated fluid flow have been demonstrated, including reversal of the flow direction. These perturbations have been associated with weld defects.

FIG. 12 SCHEMATIC SHOWING TYPICAL FLOW PATTERN IN A SUBMERGED ARC WELD POOL. SOURCE: REF 18

References cited in this section

18. J.F. LANCASTER, THE PHYSICS OF WELDING, INTERNATIONAL INSTITUTE OF WELDING, PERGAMON PRESS, OXFORD, 1984, P 243 19. F. EICHHORN AND A. ENGEL, "MASS TRANSFER IN THE WELD POOL," DOCUMENT 201-70, INTERNATIONAL INSTITUTE OF WELDING, STUDY GROUP 212, 1970 Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

References

1. N. CHRISTENSEN, V. DE L. DAVIES, AND K. GJERMUNDSEN, DISTRIBUTION OF TEMPERATURES IN ARC WELDING BR. WELD.J., VOL 12, 1965, P 54-75 2. K. ISHIZAKE, K. MURAI, AND Y. KANBE, "PENETRATION IN ARC WELDING AND CONVECTION IN MOLTEN METAL," DOCUMENT 77-66, INTERNATIONAL INSTITUTE OF WELDING, STUDY GROUP 212, 1966 3. J. THOMSON, ON CERTAIN CURIOUS MOTIONS OBSERVABLE AT THE SURFACES OF WINE AND OTHER ALCOHOLIC LIQUORS, PHILOS.MAG., VOL 10, 1855, P 330-333 4. C. MARANGONI, ÜBER DIE AUSBREITUNG DER TROPFEN EINER FLUESSIGKEIT AUF DER OBERFLAESCHE EINER ANDEREN, ANN. PHYS. CHEM., VOL 143 (NO. 7), 1871, P 337-354 5. C.R. HEIPLE AND J.R. ROPER, MECHANISM FOR MINOR ELEMENT EFFECT ON GTA FUSION ZONE GEOMETRY, WELD. J., VOL 61, 1982, P 97S-102S 6. B.J. KEENE, K.C. MILLS, AND R.F. BROOKS, SURFACE PROPERTIES OF LIQUID METALS AND THEIR EFFECTS ON WELDABILITY, MATER. SCI.TECHNOL., VOL 1, 1985, P 568-571 7. P. BURGARDT AND C.R. HEIPLE, INTERACTION BETWEEN IMPURITIES AND WELDING PARAMETERS IN DETERMINING GTA WELD SHAPE, WELD. J., VOL 65, 1986, P 150S-155S 8. M.J. TINKLER, I. GRANT, G. MIZUNO, AND C. GLUCK, WELDING 304L STAINLESS STEEL

TUBING HAVING VARIABLE PENETRATION CHARACTERISTICS, THE EFFECTS OF RESIDUAL, IMPURITY, AND MICRO-ALLOYING ELEMENTS ON WELDABILITY AND WELD PROPERTIES, P.H.M. HART, ED., PAPER 29, THE WELDING INSTITUTE, CAMBRIDGE, 1984 9. T. ZACHARIA, S.A. DAVID, J.M. VITEK, AND T. DEBROY, WELD POOL DEVELOPMENT DURING GTA AND LASER BEAM WELDING OF TYPE 304 STAINLESS STEEL, PART 1-THEORETICAL ANALYSIS, WELD. J., VOL 68, 1989, P 499S-509S 10. P. BURGARDT AND R.D. CAMPBELL, CHEMISTRY EFFECTS ON STAINLESS STEEL WELD PENETRATION, FERROUS ALLOY WELDMENTS, TRANS TECH PUBLICATIONS, SWITZERLAND, 1992, P 379-415 11. C.R. HEIPLE, P. BURGARDT, AND J.R. ROPER, THE EFFECT OF TRACE ELEMENTS ON GTA WELD PENETRATION, MODELING OF CASTING AND WELDING PROCESSES II, J.A. DANTZIG AND J.T. BERRY, ED., TMS-AIME, 1984, P 193-205 12. E. HALMOY, THE PRESSURE OF THE ARC ACTING ON THE WELD POOL, ARC PHYSICS AND WELD POOL BEHAVIOR, THE WELDING INSTITUTE, CAMBRIDGE, 1979, P 259-266 13. P.S. WEI AND W.H. GIEDT, SURFACE TENSION GRADIENT-DRIVEN FLOW AROUND AN ELECTRON BEAM WELDING CAVITY, WELD. J.,VOL 64, 1985, P 251S-259S 14. G.L. MARA, E.R. FUNK, R.C. MCMASTER, AND P.E. PENCE, PENETRATION MECHANISMS OF ELECTRON BEAM WELDING AND THE SPIKING PHENOMENON, WELD. J., VOL 53, 1974, P 246S-251S 15. H. TONG AND W.H. GIEDT, A DYNAMIC INTERPRETATION OF ELECTRON BEAM WELDING, WELD. J., VOL 49, 1970, P 259S-266S 16. W.G. ESSERS AND R. WALTER, HEAT TRANSFER AND PENETRATION MECHANISMS WITH GMA AND PLASMA-GMA WELDING, WELD. J., VOL 62, 1981, P 37S-42S 17. J.C. AMSON AND G.R. SALTER, ANALYSIS OF THE GAS-SHIELDED CONSUMABLE METALARC WELDING SYSTEM, BR. WELD. J., VOL 10, 1963, P 472-483 18. J.F. LANCASTER, THE PHYSICS OF WELDING, INTERNATIONAL INSTITUTE OF WELDING, PERGAMON PRESS, OXFORD, 1984, P 243 19. F. EICHHORN AND A. ENGEL, "MASS TRANSFER IN THE WELD POOL," DOCUMENT 201-70, INTERNATIONAL INSTITUTE OF WELDING, STUDY GROUP 212, 1970 Fluid Flow Phenomena During Welding C.R. Heiple and P. Burgardt, EG&G Rocky Flats

Selected References

• P. BURGARDT, WELDING VARIABLE EFFECTS ON WELD SHAPE IN ELECTRON BEAM WELDING OF STEELS, FERROUS ALLOY WELDMENTS, TRANS TECH PUBLICATIONS, SWITZERLAND, 1992, P 269-328 • P. BURGARDT AND R.D. CAMPBELL, CHEMISTRY EFFECTS ON STAINLESS STEEL WELD PENETRATION, FERROUS ALLOY WELDMENTS, TRANS TECH PUBLICATIONS, SWITZERLAND, 1992, P 379-415 • C.R. HEIPLE AND P. BURGARDT, PENETRATION IN GTA WELDING, WELDABILITY OF MATERIALS, R.A. PATTERSON AND K.W. MAHIN, ED., ASM INTERNATIONAL, 1990, P 73-80 • C.R. HEIPLE AND J.R. ROPER, THE GEOMETRY OF GAS TUNGSTEN ARC, GAS METAL ARC, AND SUBMERGED ARC WELD BEADS, WELDING: THEORY AND PRACTICE, D.L. OLSON, R.D. DIXON, AND A.L. LIBY, ED., NORTH-HOLLAND ELSEVIER SCIENCE PUBLISHERS, AMSTERDAM, 1990, 1-34 • K.C. MILLS AND B.J. KEENE, FACTORS AFFECTING VARIABLE WELD PENETRATION,

INT.MATER.REVIEW, VOL 35, 1990, P 185-216 Transfer of Heat and Mass to the Base Metal in Gas-Metal Arc Welding Herschel B. Smartt, Idaho National Engineering Laboratory

Introduction HEAT AND MASS TRANSFER in arc welding is normally studied from the standpoint of the weld pool and heataffected zone (HAZ); however, it is also instructive to examine heat and mass transfer from the arc to the base metal. This article describes the latter topic in terms of the gas-metal arc welding (GMAW) process and provides practical information related to the development of welding procedures and the general operation of the process. Welding procedures emphasize control of parameters such as electrode speed (or current), voltage, welding speed, contact tube-to-base metal distance, as well as current pulse parameters for out-of-position welding. It is therefore easy to overlook the fact that the process is simply a source of heat and mass inputs to the weldment. Melting of the base metal, dilution of the filler metal, solidification of the weld bead, microstructural development in the weld bead and HAZ, and thermomechanical distortion and residual stresses all follow from the heat and mass inputs. The conventional parameters identified above arc variables that control the heat and mass inputs. An example of the relationship between the conventional parameters of electrode speed and welding speed to heat and mass transferred to the weldment is shown in Fig. 1 for certain conditions, as applied to the welding of thick-section steel.

FIG. 1 PLOT OF WELDING SPEED VERSUS ELECTRODE SPEED AS FUNCTION OF HEAT TRANSFER PER LENGTH OF WELD, H, AND MASS TRANSFER EXPRESSED IN TERMS OF REINFORCEMENT, G. POWER SUPPLY OPENCIRCUIT VOLTAGE, E0, IS 32 V; CONTACT TUBE-TO-BASE METAL DISTANCE, CT, IS 15.9 MM (0.625 IN.). SHADED AREA DENOTES REGION IN WHICH SPRAY AND STREAMING TRANSFER MODES OCCUR; GLOBULAR TRANSFER OCCURS AT LOWER ELECTRODE SPEEDS, AND ELECTRODE CONTACTS THE WELD POOL AT HIGHER ELECTRODE SPEEDS. SOURCE: REF 1

The issues described in this article include the: • •

TOTAL HEAT TRANSFERRED TO THE BASE METAL PARTITIONING OF HEAT TRANSFER BETWEEN THE ARC AND THE MOLTEN ELECTRODE

• • •

DROPLETS TRANSFER MODES OF THE DROPLETS ROLE OF THE ARC IN DROPLET TRANSFER SIMPLE MODEL FOR WELDING PROCEDURE DEVELOPMENT BASED UNDERSTANDING OF HEAT AND MASS TRANSFER TO THE BASE METAL

ON

AN

Reference

1. H.B. SMARTT, IDAHO NATIONAL ENGINEERING LABORATORY, 1992 Transfer of Heat and Mass to the Base Metal in Gas-Metal Arc Welding Herschel B. Smartt, Idaho National Engineering Laboratory

Heat Transfer The total transfer of heat, H, (neglecting preheating) from the GMAW process to the weldment per unit time is given

by:

H = ηEI

(EQ 1)

where E is voltage, I is current, and η is the heat-transfer efficiency. The rate at which heat is transferred to the weldment per unit length of weld is given by: H=

η EI R

(EQ 2)

where R is welding speed. Calorimeter-based heat-transfer experiments reveal that the heat-transfer efficiency for welding thick-section steel is nominally 80 to 90%, as indicated in Fig. 2. The total heat-transfer efficiency is altered somewhat by changing other parameters. For example, it increases slightly as the power supply open-circuit voltage is decreased (for a silicon controlled rectifier regulated power supply and it increases slightly with increasing contact tube-to-base metal distance. However, 85% is a reasonable estimate for most conditions.

FIG. 2 PLOT OF HEAT-TRANSFER EFFICIENCY TO BASE METAL VERSUS ELECTRODE-SPEED FOR 0.89 MM (0.035 IN.) DIAMETER STEEL ELECTRODE IN AN AR-2% O2 SHIELD GAS. TOTAL HEAT-TRANSFER EFFICIENCY

IS SHOWN PARTITIONED INTO ARC AND MOLTEN DROP COMPONENTS. POWER SUPPLY OPEN-CIRCUIT VOLTAGE, E0, IS 32 V; CONTACT TUBE-TO-BASE METAL DISTANCE, CT, IS 15.9 MM (0.625 IN.). SOURCE: REF 2

Partitioning of Heat Transfer. In the GMAW process, the molten droplets of electrode material carry a significant portion of the total heat transferred to the weld pool. This is seen in calorimetry experiments (Fig. 2), where the total heattransfer efficiency of the GMAW process is partitioned into those portions associated with transfer by the arc and by the molten droplets. At low electrode speeds, about 60% of the total heat transferred is associated with the arc. As electrode speed increases, the fraction of total heat transferred associated with the droplets increases, reaching nominally 50% at current levels in excess of about 220 A (that is, at about 230 mm/s, or 9.1 in./s, electrode speed) for the conditions used.

Reference cited in this section

2. A.D. WATKINS, "HEAT TRANSFER EFFICIENCY IN GAS METAL ARC WELDING," MASTER'S THESIS, UNIVERSITY OF IDAHO, APRIL 1989 Transfer of Heat and Mass to the Base Metal in Gas-Metal Arc Welding Herschel B. Smartt, Idaho National Engineering Laboratory

Mass Transfer Droplet Transfer Modes. Although the International Institute of Welding (IIW) lists eight distinct metal transfer modes

(Ref 3), the modes commonly used in U.S. welding practice are globular, spray, streaming, rotating, and short circuiting. The mode terms "drop" and "repelled" used by the IIW are often referred to as "globular," and the mode term "projected" is generally referred to as "spray." The globular, spray, streaming, and short-circuiting transfer modes are shown in Fig. 3, 4, 5, 6.

FIG. 3 GLOBULAR TRANSFER MODE IN GAS-METAL ARC WELDING OF STEEL. (A) SCHEMATIC SHOWING TRANSFER OF ELECTRODE MATERIAL GLOBULES ONTO CATHODE BASE METAL. (B) HIGH-SPEED PHOTOGRAPH OF GLOBULAR METAL TRANSFER

FIG. 4 SPRAY TRANSFER MODE IN GAS-METAL ARC WELDING OF STEEL. (A) SCHEMATIC SHOWING TRANSFER OF ELECTRODE MATERIAL DROPLETS ONTO CATHODE BASE METAL. (B) HIGH-SPEED PHOTOGRAPH OF SPRAY METAL TRANSFER MODE

FIG. 5 HIGH-SPEED PHOTOGRAPH OF STREAMING TRANSFER MODE IN GAS-METAL ARC WELDING OF STEEL

FIG. 6 SHORT-CIRCUITING TRANSFER MODE IN GAS-METAL ARC WELDING OF STEEL. (A) SCHEMATIC SHOWING TRANSFER OF ELECTRODE MATERIAL BY SURFACE TENSION OF WELD POOL ONTO CATHODE BASE METAL. (B) HIGH-SPEED PHOTOGRAPH OF MATERIAL TRANSFER WHEN ARC LENGTH IS VERY SHORT. (C) HIGH-SPEED PHOTOGRAPH OF CYCLIC SHORTING OF ARC BY THE ELECTRODE DURING METAL TRANSFER TO WELD POOL. (D) HIGH-SPEED PHOTOGRAPH OF VIOLENT ARC REIGNITION WITH ASSOCIATED SPATTER

When all other parameters are held constant, the metal transfer mode at the lowest wire feed speed (and associated current level) is globular. As wire speed (and therefore current) increases, the mode changes rapidly from globular to spray. With an additional increase in wire speed (and current), spray transfer becomes streaming transfer. Figure 7 shows the droplet sizes during the transition from globular to spray to streaming transfer for direct and pulsed current welding. If an adequately high current, contact tube-to-base metal distance, and voltage exist, then rotating transfer, wherein the lower part of the electrode becomes molten over a considerable length and rotates in a helical spiral under the influence of the magnetic field surrounding the arc, can occur. As it rotates, a controlled stream of droplets is transferred from the electrode tube to the weld pool over a relatively wide area. Additional increases in wire feed/current at low voltage shorten the arc length and, eventually, the wire stubs into the weld pool. In addition, with appropriate conditions (for example, carbon dioxide, argon-carbon dioxide mixtures, and helium-based shielding gases), droplets can be transferred directly, by surface tension forces, after contact of the drop with the weld pool, a condition called short-circuiting transfer.

FIG. 7 DROPLET SIZES ATTAINABLE UNDER SELECTED CURRENT PULSING CONDITIONS IN GAS-METAL ARC WELDING OF STEEL. SOURCE: REF 4

In practical applications, the optimum transfer mode depends in part on the thickness of the base metal being welded. For example, very thin sections (in all positions) require the short-circuiting mode (with low current levels and appropriate settings of voltage and other operating parameters, including shielding gas composition). Thicker sections show best results with spray or streaming transfer. These transfer modes also produce high heat input, maximum penetration, and a high deposition rate. In welding steel, they are generally limited to welding that occurs in the flat position and the horizontal fillet position, except when pulsed current is used. Rotating transfer can be used in a deep groove in thicksection material. One- and two-pass heavy fillet welds are also a major area of application for this process variation. Globular transfer (Fig. 3) involves a droplet that, generally, is much larger in diameter than the electrode wire. Globular transfer can involve a transfer rate of about 1 to 10 drops/s, and the arc has a soft, rounded appearance. Droplet detachment and transfer are mainly due to the gravitational force, which limits globular transfer to in-position welding. For spray transfer (Fig. 4), droplet and electrode diameters are roughly equivalent. Spray transfer can involve 100 drops/s, and the arc noticeably contracts, or "stiffens." Drops usually travel in-line down the center of the arc, but several drops may be in flight at the same instant. In streaming transfer (Fig. 5), a well-developed liquid column extends from the solid electrode down into the arc and breaks into small droplets before contacting the weld pool (see the discussion on spatter below). Streaming transfer can involve 1000 drops/s, and the arc has a characteristic "cone" shape, as shown in Fig. 5. In short-circuiting transfer, the arc is very short (Fig. 6a). During metal transfer, the undetached molten droplet contacts the weld pool, shorting out the arc (Fig. 6b), which then extinguishes. Surface tension plays an important role in transferring the drop to the weld pool, but detachment of the drop from the electrode is due to electromagnetic pinch forces. Reignition of the arc is violent (Fig. 6c), resulting in considerable scatter. As seen in Fig. 7, the transition from one mode to another actually involves continuous variations in droplet size. Because the product of droplet volume and transfer rate equals the electrode melting rate (in appropriate dimensions), there is also a continuous variation in droplet transfer rate. Detailed experimental studies have shown that rapid, cyclic transition from globular to spray to globular, and so on, occurs in the transition region between globular and spray transfer (Ref 5).

Although it is generally not possible for the welder to see droplet transfer events in the spray and streaming transfer modes (except under ideal conditions for spray transfer), the arc shape changes and associated changes in electrical and acoustic noise allow a trained welder to readily identify all of the transfer modes. The section "Electrical and Acoustic Signals" in this article describes this event more fully. Weld Reinforcement. Assuming that mass is conserved as welding occurs, the transverse cross-sectional area of the

weld bead added to the weldment for a single weld pass, which is called the reinforcement, G, is given by: G=

S πd² R 4

(EQ 3)

where S is electrode speed, R is welding speed, and d is electrode diameter. For the spray and streaming transfer modes, the assumption that mass is conserved is reasonably good. However, because short circuiting and other conditions generate considerably spatter, the actual reinforcement will be slightly less than the value calculated by Eq 3. Droplet Velocity and Temperature. In spray and streaming transfer, the droplets are accelerated rapidly through the

arc to the weld pool. Velocities of 1 m/s (40 in./s) are typical, increasing with voltage (Fig. 8). Calorimetry-based experimental results indicate that steel droplets can reach temperatures of approximately 2600 K (4220 °F) (Ref 2). This explains why the droplets transport one-half of the total heat transferred to the base metal.

FIG. 8 DROPLET VELOCITIES FOR SPRAY TRANSFER IN GAS-METAL ARC WELDING OF STEEL AT THREE DIFFERENT OPEN-CIRCUIT VOLTAGES. SOURCE: REF 5

A result of the activity of molten electrode droplets in heat transfer to the base metal is that they also play an important role in convective heat transport in the weld pool and, thus, in weld penetration (Ref 6). This can be seen in Fig. 9, a transverse cross section of a gas-metal arc bead-on-plate weld on carbon steel. The region of deep penetration in the center of the weld bead is associated with the heat convected to the lower portion of the weld pool by the entering droplets.

FIG. 9 TRANSVERSE CROSS SECTION OF GAS-METAL ARC BEAD-ON-PLATE WELD IN CARBON STEEL TO SHOW DEEP PENETRATION IN THE WELD BEAD CENTER GENERATED BY MOLTEN ELECTRODE DROPLETS

Electrical and Acoustic Signals. As the droplets of molten metal detach from the electrode, an almost instantaneous change occurs in the electrode extension, which results in a sudden, although small, change in the electrical resistance between the contact tube and the base metal. The results are spikes in the secondary circuit voltage and welding current, accompanied by pressure changes in the arc. Thus, both electrical and acoustic noise (Fig. 10) are generated by the GMAW process and are characteristic of the droplet transfer mode. The waveforms for globular transfer show prominent spikes associated with individual droplet transfer events. This can also be seen in Fig. 11, a plot of the computer-digitized current for a weld on carbon steel, where a change from globular to spray transfer has occurred at about 8 s into the weld. The power spectra of the current changes dramatically for the two transfer modes (Fig. 12), leading to a means of detecting transfer mode during welding.

FIG. 10 AUDIO, CURRENT, AND VOLTAGE DURING GLOBULAR TRANSFER IN GAS-METAL ARC WELDING OF STEEL. SOURCE: REF 7

FIG. 11 PLOT OF DIGITIZED CURRENT VERSUS TIME DURING A GAS-METAL ARC WELD IN CARBON STEEL IN WHICH A TRANSITION FROM GLOBULAR TO SPRAY TRANSFER OCCURRED (AT ~T = 8 S) AS CURRENT WAS INCREASED. SOURCE: REF 1

FIG. 12 WELDING CURRENT POWER SPECTRA FOR GLOBULAR TRANSFER AND SPRAY TRANSFER MODES DURING GAS-METAL ARC WELDING OF STEEL. SOURCE: REF 1

Several specific techniques have been identified (Ref 8) for detecting the mode of metal transfer: Fourier transform, standard deviation, peak ratios, and integrated amplitude of the current and voltage signal. In a similar manner, spatter, lack of shielding gas, and contact tube wear can also be detected. Similar power spectra changes also occur in the secondary circuit voltage and acoustic noise. The change in acoustic noise allows the aural detection of the transfer mode. An experienced welder can readily hear, for example, the difference between globular and spray transfer. Contact of the molten droplet with the weld pool, while the droplet is still attached to the electrode, also results in a sudden change in electrical resistance. This is important, because spatter is generally produced when this electrode-to-

weld pool contact breaks. (Spatter can also be produced by rapid expansion of gas bubbles in the electrode as it melts.) The resulting electrical and acoustic noise can be used to detect that spatter is being produced.

References cited in this section

1. H.B. SMARTT, IDAHO NATIONAL ENGINEERING LABORATORY, 1992 2. A.D. WATKINS, "HEAT TRANSFER EFFICIENCY IN GAS METAL ARC WELDING," MASTER'S THESIS, UNIVERSITY OF IDAHO, APRIL 1989 3. J.F. LANCASTER, THE PHYSICS OF WELDING, 2ND ED., INTERNATIONAL INSTITUTE OF WELDING, PERGAMON PRESS, 1986, P 233 4. T.W. EAGAR, D.E. HARDT, H.B. SMARTT, AND J.A. JOHNSON, PROCESS CONTROLLABILITY IN GAS METAL ARC WELDING: MICRO AND MACROSCOPIC PROBLEMS, PROC. EIGHTH SYMPOSIUM ON ENERGY SCIENCES, CONF-9005183, U.S. DEPARTMENT OF ENERGY, ARGONNE NATIONAL LABORATORY, MAY 1990, P 51-67 5. D.E. CLARK, C. BUHRMASTER, AND H.B. SMARTT, DROPLET TRANSFER MECHANISMS IN GMAW, PROC. 2ND INTERNATIONAL CONFERENCE ON TRENDS IN WELDING RESEARCH (GATLINBURG, TN), 1989 6. W.G. ESSERS AND R. WALTER, HEAT TRANSFER AND PENETRATION MECHANISMS WITH GMA AND PLASMA-GMA WELDING, WELD.J., VOL 60 (NO. 2), 1981, P 37-S TO 42-S 7. J.A. JOHNSON, N.M. CARLSON, AND H.B. SMARTT, DETECTION OF METAL TRANSFER MODE IN GMAW, PROC. 2ND INTERNATIONAL CONFERENCE ON TRENDS IN WELDING RESEARCH (GATLINBURG, TN), 1989 8. G. ADAM AND T.A. SIEWERT, SENSING OF GMAW DROPLET TRANSFER MODES USING AN ER100S-1 ELECTRODE, WELD.J., VOL 69 (NO. 3), MARCH 1990, P 103-S TO 108-S Transfer of Heat and Mass to the Base Metal in Gas-Metal Arc Welding Herschel B. Smartt, Idaho National Engineering Laboratory

Procedure Development The preceding discussion leads to a logical approach toward selecting welding parameters during the development of welding procedures. First, given a nominal power supply open-circuit voltage and contact tube-to-base metal distance, the electrode speed can be set to determine the welding current. It can be shown that: S=

4η GI ( E0 + HI ) Hπ d ²

(EQ 4)

where E0 is the power supply open-circuit voltage. It is assumed that the proper selection of electrode type and shielding gas have been made, and that the electrode diameter is appropriate for the application. Second, the open-circuit voltage is adjusted to give either an acceptable arc length or electrode extension. The objective is to prevent either burnback (transfer of the arc to the contact tube) or spatter caused by shorting of the molten electrode droplets with the weld pool prior to detachment from the electrode. Fine adjustment of voltage may be necessary to obtain a clean start of the process, thus avoiding a short period of spatter or globular transfer following arc ignition. The combination of current and voltage for a given contact tube-to-base metal distance determines the droplet transfer mode. Third, the welding current determines the melting rate of the electrode. The electrode melting rate (Mrp) is a quadratic function of current (Ref 9):

M rp = (C1 + C2 A) I + C3

LI ² A²

(EQ 5)

where A is the cross-sectional area of the electrode and C1, C2, C3, and i are constants. Another researcher (Ref 10) defines Mrp by: M rp = C4 I + C5

LI ² A

(EQ 6)

where C4 and C5 are constants. Given a desired melting rate, the welding travel speed is set to obtain the desired weld bead reinforcement. The current, voltage, and weld travel speed have all now been set, thus determining the heat input per length of weld (H = ηEI/R). It should be noted that it is actually possible to independently vary heat input and mass input (in terms of weld bead reinforcement) to the weld, over at least a small range. This is shown in Fig. 13, where the same data presented in Fig. 1 are replotted in terms of reinforcement as a function of heat input per length of weld.

FIG. 13 PLOT OF REINFORCEMENT VERSUS HEAT INPUT TO SHOW PARAMETERS THAT FAVOR SPRAY TRANSFER MODE FOR GAS-METAL ARC WELDING OF STEEL. POWER SUPPLY OPEN-CIRCUIT VOLTAGE, E0, IS 32 V; CONTACT TUBE-TO-BASE METAL DISTANCE, CT, IS 15.9 MM (0.625 IN.).

The above discussion on procedure development ignores second-order effects, such as the dependence of arc length on weld travel speed. It should also be realized that several iterations through the steps defined above may be required for final parameter determination. Proper power supply settings must be used, and code requirements must be met.

References cited in this section

9. A. LESNEWICH, CONTROL OF MELTING RATE AND METAL TRANSFER IN GAS-SHIELDED

METAL-ARC WELDING, PART II--CONTROL OF METAL TRANSFER, WELD.J., VOL 37 (NO. 9), SEPT 1958, P 418-S TO 425-S 10. J.H. WASZINK AND G.J.P.M. VAN DEN HEUVEL, HEAT GENERATION AND HEAT FLOW IN THE FILLER METAL IN GMA WELDING, WELD.J., VOL 61 (NO. 8), AUG 1982, P 269-S TO 282-S Transfer of Heat and Mass to the Base Metal in Gas-Metal Arc Welding Herschel B. Smartt, Idaho National Engineering Laboratory

References

1. H.B. SMARTT, IDAHO NATIONAL ENGINEERING LABORATORY, 1992 2. A.D. WATKINS, "HEAT TRANSFER EFFICIENCY IN GAS METAL ARC WELDING," MASTER'S THESIS, UNIVERSITY OF IDAHO, APRIL 1989 3. J.F. LANCASTER, THE PHYSICS OF WELDING, 2ND ED., INTERNATIONAL INSTITUTE OF WELDING, PERGAMON PRESS, 1986, P 233 4. T.W. EAGAR, D.E. HARDT, H.B. SMARTT, AND J.A. JOHNSON, PROCESS CONTROLLABILITY IN GAS METAL ARC WELDING: MICRO AND MACROSCOPIC PROBLEMS, PROC. EIGHTH SYMPOSIUM ON ENERGY SCIENCES, CONF-9005183, U.S. DEPARTMENT OF ENERGY, ARGONNE NATIONAL LABORATORY, MAY 1990, P 51-67 5. D.E. CLARK, C. BUHRMASTER, AND H.B. SMARTT, DROPLET TRANSFER MECHANISMS IN GMAW, PROC. 2ND INTERNATIONAL CONFERENCE ON TRENDS IN WELDING RESEARCH (GATLINBURG, TN), 1989 6. W.G. ESSERS AND R. WALTER, HEAT TRANSFER AND PENETRATION MECHANISMS WITH GMA AND PLASMA-GMA WELDING, WELD.J., VOL 60 (NO. 2), 1981, P 37-S TO 42-S 7. J.A. JOHNSON, N.M. CARLSON, AND H.B. SMARTT, DETECTION OF METAL TRANSFER MODE IN GMAW, PROC. 2ND INTERNATIONAL CONFERENCE ON TRENDS IN WELDING RESEARCH (GATLINBURG, TN), 1989 8. G. ADAM AND T.A. SIEWERT, SENSING OF GMAW DROPLET TRANSFER MODES USING AN ER100S-1 ELECTRODE, WELD.J., VOL 69 (NO. 3), MARCH 1990, P 103-S TO 108-S 9. A. LESNEWICH, CONTROL OF MELTING RATE AND METAL TRANSFER IN GAS-SHIELDED METAL-ARC WELDING, PART II--CONTROL OF METAL TRANSFER, WELD.J., VOL 37 (NO. 9), SEPT 1958, P 418-S TO 425-S 10. J.H. WASZINK AND G.J.P.M. VAN DEN HEUVEL, HEAT GENERATION AND HEAT FLOW IN THE FILLER METAL IN GMA WELDING, WELD.J., VOL 61 (NO. 8), AUG 1982, P 269-S TO 282-S Arc Physics of Gas-Tungsten Arc Welding J.F. Key, EG&G Idaho, Inc.

Introduction THE GAS-TUNGSTEN ARC WELDING (GTAW) process is performed using a welding arc between a nonconsumable tungsten-base electrode and the workpieces to be joined. C.E. Jackson defined a welding arc as "a sustained electrical discharge through a high-temperature conducting plasma producing sufficient thermal energy so as to be useful for the joining of metals by fusion." This definition is a good foundation for the discussion that follows.

The physics of GTAW are fundamental to all arc processes and are more straightforward, because the complications of materials (for example, filler and flux) transferred through and interacting with the arc can be avoided. Geometrically, the arc discharge in GTAW is between a rod-shaped tungsten electrode and a planar-shaped electrode, that is, the workpiece. Pure tungsten electrodes are less expensive and, possibly, more environmentally compatible than those with rare earth or other oxide additions. They are used for lower-specification welds, where tungsten contamination that is caused by the molten electrode surface can be tolerated. They are also used for alternating current (ac) welding of aluminum, copper, magnesium, and thin sections of low-alloy and stainless steels. Analysis of the arc discharge is separated into electrode regions and the arc column. The electrode regions are confined to very small distances from the electrode surfaces, have very high electrical and thermal fields, and have much higher current density, because of the contraction of the arc to a small spot. As a result, electrode regions for both the cathode and the anode are difficult to analyze by diagnostic measurements and theoretical computation. This situation must be remedied for a thorough understanding of the process, because the process parameters control the arc discharge at the cathode, with the anode serving as the connection to ground. The arc column, on the other hand, is relatively easy to analyze, but is important primarily as a means to deduce arc characteristics at the electrodes. Polarity. The GTAW process generally utilizes a direct current (dc) arc, where the tungsten electrode has a negative polarity. The tungsten electrode thus becomes the cathode and the workpiece becomes the anode. The polarity is called straight polarity, or direct current electrode negative (DCEN).

Reverse polarity, or direct current electrode positive (DCEP), is literally the reverse of DCEN. The workpiece is the cathode and the tungsten electrode serves as the anode. Because most heat is generated at the anode in the GTAW process, DCEP is used for welding certain thin-section, low melting point materials when DCEN would be likely to cause excessive penetration or burn-through. Either alternating current or DCEP is used for removing an oxide film from the surface of the weld pool or workpiece. The oxide film promotes emission during the half-cycle (ac) when the workpiece is negative polarity. As the oxide is depleted, the emission moves to a new location that has a high enough oxide content to sustain the discharge of electrons. The arc root or cathode spot where the emission occurs is highly mobile in ac or DCEP and, as a result, the arc is much less stable than in DCEN. Gas Shielding. In all cases, the arc and both electrodes are shielding by gas, usually an inert gas or a gas mixture. Argon

and argon-helium mixtures are used most often, although argon-hydrogen mixtures are used for some applications. The GTAW process may simply utilize the arc to fuse the workpieces together without the addition of filler materials (autogenous) or filler may be added to the molten pool to fill grooves in thicker weldments. A reasonable understanding of welding arc fundamentals and the GTAW process requires a more thorough discussion of the electrode regions of the arc and the arc column. Arc Physics of Gas-Tungsten Arc Welding J.F. Key, EG&G Idaho, Inc.

Electrode Regions and Arc Column The cathode and anode are similar in several respects. Both exhibit a voltage drop caused by a space charge that covers a very thin region over their surfaces, and the arc is significantly contracted on the surfaces. Figure 1 shows that the total arc voltage is partitioned between the electrode drops and arc column. The relative magnitude of these drops depends on welding parameters and electrode material.

FIG. 1 PLOT OF RELATIVE ARC VOLTAGE DISTRIBUTION VERSUS RELATIVE ARC LENGTH BETWEEN ELECTRODES

The arc discharge requires a flow of electrons from the cathode through the arc column to the anode, regardless of polarity or whether ac or dc is used. Two cases of electron discharge at the cathode will be discussed: thermionic emission and nonthermionic emission, also called cold cathode, or field emission. Thermionic emission results from joule heating (resistance) of the cathode by the imposed welding current until the electron energy at the cathode tip exceeds the work function (energy required to strip off an electron). This case applies to the general case of DCEN, where the tungsten electrode is the emitter, or cathode. Pure tungsten electrodes have to be heated to their melting point to achieve thermionic emission. Once molten, the equilibrium tip shape becomes a hemisphere, and a stable arc results from uniform emission over this surface. Thoria (ThO2), zirconia (ZrO2), or ceria (CeO2) are added to pure tungsten in amounts up to 2.2 wt% ThO2, 0.4 wt% ZrO2, or 3.0 wt% CeO2 to lower the work function, which results in thermionic emission at lower temperatures and avoids melting the cathode tip. These electrodes typically have a ground conical tip, and thermionic emission is localized to a cathode spot. Thermionic emission creates a cloud of electrons, called a space charge, around the cathode. If a second electrode at a higher potential is nearby (the workpiece, in this case), then the electrons will flow to it, thus establishing the arc. Nonthermionic, or field, emission creates an electron discharge with a very high electric field, typically exceeding 109 V/m. This intense electric field literary pulls electrons out of a relatively cold or unheated cathode. This would not appear to be applicable to welding until one considers that for reverse polarity or DCEP, a condensation of positive ions from the arc column can build up in a very thin (1 nm, or 0.04 μin.) layer over the cathode surface, creating a very high localized electric field even though the cathode voltage drop may only amount to several volts. An oxide layer, which is always present on the cathode surface in an actual weld, facilitates the discharge with a source of lower work function electrons. When the oxide layer is very thin (on the order of one atom layer), emission occurs via a tunneling mechanism through the film to an emitting site. Thicker oxide films exhibit locally conducting spots at the end of filamentary channels through the oxide. Large currents flow in these channels, which are on the order of 100 nm (4 μin.) in diameter and have lifetimes of 0.001 to 1 μs. The cathode cleaning action, which is one of the principal reasons to use DCEP or ac, results from stripping away the oxide film at the emitting sites by very small and intense jets of metal vapor and debris. It becomes obvious that a practical implication of the short lifetime of these cathode spots is a generally unstable arc that is due to the necessity of continual movement of the cathode spot to undepleted regions of oxide film. Arc instability is undesirable and DCEP or ac is only used when cathodic cleaning or the minimizing of heat input to the workpiece is a higher priority than optimizing weld bead shape and location. Anode. Welding process parameters (for example, current and voltage) control the arc discharge at the cathode. Although

the electron flow enters the anode through the anode spot and constitutes 85% of the energy going into the weld pool, thus making current density the single most important welding parameter that determines pool shape, events at the anode can

only be controlled indirectly by controlling the cathode. Anode spot stability does depend somewhat on shielding gas composition and the shape of the anode (that is, weld groove). Current density and heat input measurements at the anode have been made to better understand how process parameters that are largely controlling events at the cathode will, in turn, influence the shape and melting rate of the weld pool. The relative contributions of heat transfer to the workpiece, in terms of the GTAW process, are shown in Fig. 2.

FIG. 2 RELATIVE HEAT TRANSFER CONTRIBUTIONS TO WORKPIECE WITH GTAW. (A) CONTRIBUTION OF INDIVIDUAL PARAMETERS TO ANODE HEAT INPUT. (B) HEAT OUTPUT AT CATHODE (WORKPIECE) RELATIVE TO WELD POOL HEAT LOSS

The Thomson effect represents the energy lost by electrons as they move from higher to lower temperatures. The sum of work function, Thomson effect, and anode fall gives an electron contribution to heat transfer of approximately 84%. The remaining 16% is due to thermal effects (that is, conduction, convection, and radiation). There are small heat losses from the pool that are due to evaporation of metal ions and radiation. Figure 3 shows that the energy distribution for this particular case approaches a Gaussian distribution (that is, normal distribution curve). High helium contents in the shielding gas have produced data that are typically better fit by a Lorentzian curve, indicating a narrower current density distribution (that is, a more-contracted arc at the anode spot).

FIG. 3 PLOT OF ELECTRON AND THERMAL CONTRIBUTIONS TO HEAT TRANSFER. A, TOTAL ARC POWER (STANDARD DEVIATION, σ , OF 0.8 MM, OR 0.031 IN.); B, ELECTRON CONTRIBUTION (σ = 0.7 MM OR 0.028 IN.); C, THERMAL CONTRIBUTION. WELD PARAMETERS: CURRENT, 10 A; VOLTAGE, 10 V; TIME, 10 S; SHIELDING GAS, ARGON; ELECTRODE ANGLE, 30°

Arc efficiency, in addition to those variables that have an effect on it, is an extremely important term in the heat transfer analysis of welding. It gives the percentage of heat dissipated in the arc that actually is captured by the workpieces and is available for melting. Arc efficiency, as a function of all GTAW welding parameters and many materials, has been determined experimentally and found to be nominally 75%. The variables having the greatest effect on arc efficiency are arc voltage and anode material. For those variables, the effect is usually no more than ±5%. Other parameters have a negligible effect. Arc Column. The electron discharge between the electrodes partially ionizes the shielding gas in its path, thus making

the arc column a conductor, or plasma. Overall, the arc column is neutral and is composed of electrons, positive gas and possibly metal ions, and neutral gas atoms. Ironically, fundamental measurements of arc properties are most easily made in the arc column, although the actual effect of these properties on the electrode region of an actual weld must still be inferred. Nevertheless, it is useful to understand fundamentals of the arc that relate essential welding variables (for example, current, voltage, electrode gap, choice of shielding gas, and electrode shape) to arc temperature, current density distribution, and gas flow structure at the anode surface. Effect of Cathode Tip Shape. For the general case of straight polarity, DCEN, the tip of the tungsten alloy cathode is

ground to a point and then truncated somewhat to prevent the sharp tip from burning off and contaminating the weld. The included angle of the cone and the diameter of the truncation under some welding conditions have a significant effect on weld pool shape. Figure 4 shows examples of the effect of these two parameters on weld pool shape.

FIG. 4 FUSION ZONE PROFILE FOR SPOT-ON-PLATE WELDS AS A FUNCTION OF ELECTRODE TIP GEOMETRY USING 100% AR AS A SHIELDING GAS. WELD PARAMETERS: CURRENT, 150 A; DURATION, 2 S

For a stationary spot-on-plate weld shielded by pure argon, the weld depth-to-width ratio increased with an increasing vertex angle up to 90° and with an increasing truncation diameter. The arc became less "bell shaped" and more "ball shaped" as the vertex angle or truncation diameter increased. These results should be a valid indication of the effect of cathode tip shape for pulsed current welding, which produces a series of overlapping spot welds. A study of bead-on-plate welds (Fig. 5) made with constant current and velocity indicated a similar but less pronounced trend. These conditions produce a tear-drop molten pool shape when viewed from above, compared to circular shape for spot or pulsed current welding. Fluid and heat flow within the pool is less uniform front-to-rear in a tear-drop-shaped pool and probably has a greater influence on pool shape than electrode tip shape.

FIG. 5 FUSION ZONE PROFILE FOR BEAD-ON-PLATE WELDS AS A FUNCTION OF ELECTRODE TIP GEOMETRY USING 100% AR AS A SHIELDING GAS. WELD PARAMETERS: CURRENT, 150 A; WELDING SPEED, 3 MM/S (0.12 IN./S)

When the arc is used in a weld groove, the relative shapes of the cathode tip and the anode groove become more important. The arc discharge from the cathode will seek a path to ground with the lowest electrical resistance. For a stable arc properly centered in the groove (for example, a root pass), the shortest path to ground should be between the cathode tip and the bottom of the groove (Fig. 6). This will require the cathode vertex angle to be somewhat less than the included angle of the groove and/or a sufficiently wide groove to ensure that the shortest path to ground is from the cathode tip to the groove bottom and not, for example, from the electrode shoulder to the groove wall, as the case would be with a 90° electrode in a 10° narrow groove. Welding in a groove places a higher priority on arc stability and location than on maximum penetration.

FIG. 6 EFFECTS OF ELECTRODE TIP GEOMETRY ON THE PATH LENGTH TO GROUND IN WELD GROOVES OF VARIOUS SHAPES. (A) 75° V-GROOVE. (B) 40° U-GROOVE. (C) 10° NARROW GROOVE

To understand the effects of tip shape, temperature distributions in the plasma were measured. Figure 7 shows that as the cathode vertex angle increases, the plasma radius of the arc column increases at mid-gap and becomes more constricted near the anode. The quantitative interpretation of these results requires theoretical modeling, which has yet to be completed.

FIG. 7 EFFECT OF VERTEX ANGLE ON GTAW ARC COLUMN TEMPERATURE DISTRIBUTION WITH 100% AR USED AS SHIELDING GAS. (A) 30° ELECTRODE VERTEX ANGLE. (B) 90° ELECTRODE VERTEX ANGLE. WELDING CURRENT, 150 A

Effect of Shielding Gas Composition. The GTAW process typically uses either an inert gas, such as argon, or an inert gas mixture, such as argon and helium, to shield the arc and the weld from atmospheric contamination. Occasionally, a slightly reactive gas mixture, such as argon with up to 15 vol% H2, is used. (The 15 vol% limit is based on safety considerations.) Shielding gas composition has a rather strong effect on arc temperature distribution and, under certain conditions, a significant effect on weld pool shape.

Figure 8 shows how shielding gas affects arc voltage. The curves would all be displaced downward for shorter arc lengths, but the relative positions would be maintained. Figure 9 shows the effects of both cathode vertex angle and shielding gas composition on weld pool shape for spot-on-plate welds. Increasing additions of helium to argon show a remarkable increase in penetration when using a 30° vertex angle. However, the effect is much less evident when using a 90° vertex angle.

FIG. 8 PLOT OF ARC VOLTAGE VERSUS ARC CURRENT FOR SELECTED INERT SHIELDING GASES. WELDING PARAMETERS: ANODE, TITANIUM; CATHODE, TUNGSTEN; POLARITY, DCEN; ARC LENGTH, 12.7 MM (0.050 IN.)

FIG. 9 EFFECT OF ELECTRODE TIP GEOMETRY AND SHIELDING GAS COMPOSITION ON WELD POOL SHAPE FOR SPOT-ON-PLATE WELDS. WELDING PARAMETERS: CURRENT, 150 A; DURATION, 2 S

To understand these phenomena, arc temperature distributions for a variety of shielding gases and mixtures, electrode shapes, current, arc voltages (electrode gaps), and anode materials have been measured in order to clarify welding arc fundamentals. Welding arcs are composed of electrons, positive gas ions, and neutral gas atoms. Some measurement techniques give the temperature of one species (electrons), whereas other techniques give the temperature of another species (neutral atoms). If the arc is in local thermodynamic equilibrium (LTE), all techniques should give the same temperature. Although the assumption of LTE used to be considered completely valid, contemporary investigations have suggested that this is not always the case and that some of the older measurements may be somewhat in error. Absolute values of arc temperature are only needed to establish boundary conditions for detailed arc modeling of temperature-dependent properties. What is of more importance to the welding engineer or technologist is the relative effect of essential variables

on heat input to the workpiece. Arc temperature measurement is one useful indication of how these variables affect the arc. When compared to Fig. 7(a), Fig. 10 shows that large additions of helium to argon decrease peak temperatures slightly and increase the plasma diameter in the plasma column. The arc appears to become a broader and more isothermal heat source. The lower peak temperature is reasonable, because a combination of the high ionization potential of helium and relatively low currents of welding arcs gives an arc column that is only slightly ionized. Figure 7 showed that a large vertex angle had a similar, but less pronounced, effect on arc temperature when adding helium to the shielding gas (that is, the axial peak temperature decreased and the plasma diameter increased).

FIG. 10 PLOT OF GTAW ARC COLUMN TEMPERATURE DISTRIBUTION AS A FUNCTION OF ANODE DISTANCE AND ARC POSITION. WELDING PARAMETERS: ELECTRODE VERTEX ANGLE, 30°; CURRENT, 150 A; SHIELDING GAS, 10AR-90HE

Figure 11 shows that doubling the current from 150 to 300 A produces an increase in plasma diameter (that is, that portion of the arc above approximately 8000 K, or 13,900 °F). This effect occurs regardless of the shielding gas composition.

FIG. 11 PLOT OF GTAW ARC COLUMN TEMPERATURE DISTRIBUTION RELATIVE TO ANODE DISTANCE AND ARC POSITION. WELDING PARAMETERS: ELECTRODE VERTEX ANGLE, 30°; CURRENT, 300 A; SHIELDING GAS, 100% AR

The arc length or gap between the electrodes is yet another process parameter that must be considered, especially for mechanized welding, where it can be kept reasonably constant. Because arc length is proportional to arc voltage, longer arcs have higher arc voltages and consume more energy for the same current. However, this increased energy is generally lost through radiation to the environment surrounding the weld and does not effectively supply additional heat to the workpieces. Mechanized welding generally utilizes rather short arc lengths (2 to 3 mm, or 0.08 to 0.12 in.), whereas manual welding uses a longer arc length. Unfortunately, relationships that establish a direct correlation between the temperature distribution in the arc column and the weld pool shape still have not been established, primarily because weld pool shape depends on other factors, such as compositionally dependent molten metal properties. Flow Structure. Shielding gas is used to displace reactive gases in the atmosphere from the vicinity of the weld. Inert

gases are preferred for the GTAW process, because they minimize unwanted gas-metal reactions with the workpieces. A uniform laminar flow of gas from the gas cup would be ideal and, in fact, is usually achieved as long as there is no welding arc. However, the arc discharge rapidly heats the gas in the arc column, and thermal expansion causes plasma jets. The lower temperature near the cathode tip in Fig. 11 is an indication of a jet pumping in cooler gas. This becomes an important factor at high currents, because these jets can depress the surface of the weld pool and alter heat transfer to it. The rapid gas expansion can cause the flow to deviate from laminar and, in extreme cases, the flow can become turbulent. Turbulence tends to mix atmospheric contaminants into the arc, often where they can do the most harm: at the surface of the molten weld pool. Holographic interferometry and Schlieren photography have been used to characterize gas flow. Figures 12, 13, and 14 show examples of laminar and turbulent flow. The flow from a commercial-design gas cup for three current levels is shown in Fig. 12. Laminar flow occurs where the fringes are generally straight and parallel. Turbulent flow is indicated by very wavy or circular fringes. Increasing current tends to make the laminar region somewhat broader and thicker, effectively increasing the area shielded from atmospheric contamination.

FIG. 12 EFFECT OF GEOMETRY ON COMMERCIAL GAS CUP LAMINAR AND TURBULENT FLOW AS DETECTED BY REAL-TIME HOLOGRAPHIC INTERFEROMETRY

FIG. 13 EFFECT OF GEOMETRY ON CONVERGING CONE CUP LAMINAR AND TURBULENT FLOW AS DETECTED BY REAL-TIME HOLOGRAPHIC INTERFEROMETRY

FIG. 14 EFFECT OF GEOMETRY ON VENTURI GAS CUP LAMINAR AND TURBULENT FLOW AS DETECTED BY REAL-TIME HOLOGRAPHIC INTERFEROMETRY

Figures 13 and 14 result from experiments with gas cup shapes that were designed to improve shielding. Figure 13 shows that a converging cone would be a very poor choice for the GTAW process, as indicated by the very small area of laminar flow and the extreme turbulence in the surrounding region.

Figure 14 is a venturi shape, which provides a large laminar flow region for all current levels and excellent shielding. This design may be somewhat better than commercial designs, but weld contamination studies should be conducted to verify this possibility. Arc Physics of Gas-Tungsten Arc Welding J.F. Key, EG&G Idaho, Inc.

Selected References

• S.S. GLICKSTEIN, "ARC-WELD POOL INTERACTIONS," DEPARTMENT OF ENERGY RESEARCH AND DEVELOPMENT REPORT WAPD-TM-1429, BETTIS ATOMIC POWER LABORATORY, AUG 1978 • C.E. JACKSON, THE SCIENCE OF ARC WELDING,WELD. J., VOL 39, 1960, P 129S-140S, 177S-190S, 225S-230S • J.F. KEY, H.B. SMARTT, J.W. CHAN, AND M.E. MCILWAIN, PROCESS PARAMETER EFFECTS ON ARC PHYSICS AND HEAT FLOW IN GTAW, PROC. INT. CONF. ON WELDING TECHNOLOGY FOR ENERGY APPLICATIONS (GATLINBURG, TN), 16-19 MAY, 1982, P 179-199 • J.F. KEY, J.W. CHAN, AND M.E. MCILWAIN, PROCESS VARIABLE INFLUENCE ON ARC TEMPERATURE DISTRIBUTION, WELD. J./WELD. RES. SUPP., VOL 62 (NO. 7), P 179-S TO 184-S • J.F. LANCASTER, ED., THE PHYSICS OF WELDING, 2ND ED., INTERNATIONAL INSTITUTE OF WELDING, PERGAMON PRESS, 1986 • H.B. SMARTT, J.A. STEWART, AND C.J. EINERSON, "HEAT TRANSFER IN GAS TUNGSTEN ARC WELDING," ASM METALS/MATERIALS TECHNOLOGY SERIES, ASM INTERNATIONAL WELDING CONGRESS (TORONTO), 14-17 OCT 1985 Power Sources for Welding F. James Grist, Miller Electric Manufacturing Company, Inc.; William Farrell and Glen S. Lawrence, Ferranti Sciaky

Introduction POWER SOURCES are apparatuses that are used to supply current and voltages that are suitable for particular welding processes. This article describes power sources for arc welding, resistance welding, and electron-beam welding. Power Sources for Welding F. James Grist, Miller Electric Manufacturing Company, Inc.; William Farrell and Glen S. Lawrence, Ferranti Sciaky

Arc Welding Power Sources Arc welding requires that an electric arc be established between an electrode and the workpiece to produce the heat needed for melting the base plate. Because utility energy is not delivered at the proper voltage and current, it must be converted to the required levels by the welding power source. Arc power sources convert the customary 240 or 480 V alternating current (ac) utility power to a range from 20 to 80 V and simultaneously increase the current proportionately. Galvanic isolation is also provided from the utility line voltage to ensure operator safety. Motor- or engine-driven welding generators are wound to deliver the correct voltage and current

directly; therefore, no transformer is necessary. Actual arc voltage is established by factors other than the power source (for example, electrode type, base-metal type, shielding gas, and standoff distance). Ratings and Standards The National Electrical Manufacturers Association (NEMA) categorizes arc welding power sources into three classes on the basis of duty cycle: • • •

CLASS I: RATED OUTPUT AT 60, 80, OR 100% DUTY CYCLE CLASS II: RATED OUTPUT AT 30, 40, OR 50% DUTY CYCLE CLASS III: RATED OUTPUT AT 20% DUTY CYCLE

Duty cycle is the ratio of arc time to total time based on a 10 min averaging period. A 60% machine will deliver 6 min of arc time and 4 min off time without overheating. In Fig. 1, curve A shows a NEMA class I (60%) 300 A rated machine that is capable of a maximum 375 A at reduced duty cycle (38%) and 232 A at 100% (continuous). Curve B represents a NEMA class II (50%) 250 A machine with a continuous duty of 176 A. Curve C represents an engine-driven machine rated at 225 A and 20% duty. It does not offer output in excess of its rating, because of a horsepower limitation of the engine.

FIG. 1 SELECTED DUTY CYCLE CURVES. A, 300 A, 60% MACHINE; B, 250 A, 50% MACHINE; C, 225 A, 20% ENGINE-DRIVEN MACHINE WITH HORSEPOWER LIMITATION

NEMA requires most machines to produce a maximum of 125% of rated output current. No minimum current is specified, but 10% (of rating) is typical. Load voltage, E, for class I and II machines is defined by:

E = 20 + 0.04I

(EQ 1)

where I is the current in amperes, below 600 A. Above 600 A, the rated load voltage remains at 44 V. Power sources intended for gas-tungsten arc welding (GTAW) service are defined by:

EGTAW = 13 + 0.012I Class III machines are constant-current types limited to 20% duty cycle.

(EQ 2)

Efficiencies (that is, ratios of power output to power input) of 68% for line-frequency machines and 88% for inverters are typical. Line current draw can be reduced in certain machines by the addition of power factor correcting capacitors connected to the primary winding. Input (line) voltages shown on the manufacturer nameplate are often multivoltage by means of tap changing for 200, 230, 460, or 575 V. Designed primarily for 60 Hz power, most systems will operate at 50 Hz, with minor changes in output. Input voltages for 50 Hz systems are typically 220, 380, and 415 V. Usual service conditions are temperature that range from 0 to 40 °C (32 to 105 °F) and altitudes of up to 1000 m (3300 ft). (Altitude is a factor because of fuel-air mixtures.) The power source nameplate must include the following specifications: • • • • • • • • •

MANUFACTURER NEMA CLASS (I, II, OR III) AND DUTY CYCLE MAXIMUM OPEN-CIRCUIT VOLTAGE NUMBER OF PHASES MAXIMUM SPEED (IN REVOLUTIONS PER MINUTE) AT NO LOAD (ROTATING EQUIPMENT) RATED LOAD (IN VOLTS) RATED LOAD (IN AMPERES) FREQUENCY (OR FREQUENCIES) IN HERTZ INPUT VOLTAGE(S)

The National Electric Code (NEC) requires that a main disconnect switch be provided for each installed welding machine. Manufacturer installation instructions list the kVA and kW rating and recommend the correct fuse and conductor size. Local codes may differ from NEC requirements. Proper grounding is a necessary safety consideration. Machines equipped with high-frequency arc stabilizers must be correctly grounded to prevent electromagnetic interference (EMI) to nearby equipment. The installer certifies (to the Federal Communications Commission) that the machine has been grounded according to manufacturer instructions. Most power sources are forced-air cooled by internal fans. Thermal sensors will interrupt output if airflow is impeded. Power Source Selection Because no single power source is right for all welding situations, it is necessary to know the processes to be used before selecting the best power source. Table 1 gives the more-common processes that use constant-current and constant-voltage power sources.

TABLE 1 SELECTION OF POWER SOURCE RELATIVE TO WELDING PROCESS TO BE EMPLOYED

WELDING PROCESS

POWER SOURCE CONSTANT CONSTANT CURRENT VOLTAGE SHIELDED METAL ARC WELDING (SMAW) • ... GAS-TUNGSTEN ARC WELDING (GTAW) • ... GAS-METAL ARC WELDING (GMAW) ... • PULSED GAS-METAL ARC WELDING (GMAW-P) • ... FLUX-CORED ARC WELDING (FCAW) ... • SUBMERGED ARC WELDING (SAW) ... • ELECTROGAS WELDING (EGW) • • ELECTROSLAG WELDING (ESW) ... • PLASMA ARC WELDING (PAW) • ...

PLASMA ARC CUTTING (PAC)

•

...

Alternating current power sources for shielded metal arc welding (SMAW) and submerged arc welding (SAW) can be as simple as a single transformer (Fig. 2a). Low-voltage machines with open-circuit voltages of 50 V or less are transformer types. Output alternating current reflects the sinusoidal waveform of the utility input line (Fig. 3a).

FIG. 2 SELECTED ALTERNATING CURRENT POWER SOURCES. (A) ALTERNATING CURRENT WELDING TRANSFORMER, ADJUSTABLE CORE OR WINDINGS. (B) THREE-PHASE SCR-CONTROLLED DC WELDING POWER SOURCE. (C) INVERTER BLOCK DIAGRAM. (D) MOTOR-GENERATOR SET. (E) ENGINE-DRIVEN ALTERNATOR WITH DC OUTPUT. (F) SECONDARY SWITCHING POWER SOURCE

FIG. 3 TYPICAL AC WELDING CURRENT WAVEFORMS. (A) SINE WAVE ALTERNATING CURRENT, SIMPLE TRANSFORMER WELDER. (B) MODIFIED SINE WAVE OBTAINED FROM MAGNETIC AMPLIFIER-TYPE WELDER. (C) SQUARE WAVE, LINE FREQUENCY, AND BALANCED DWELL. (D) SQUARE WAVE WITH UNBALANCED DWELLS, EP VERSUS EN. (E) VARIABLE-POLARITY AC WITH DWELL, AMPLITUDE, AND FREQUENCY ALL ADJUSTABLE

Magnetically controlled equipment (that is, a saturable reactor or magnetic amplifier) used for GTAW delivers a modified sinewave output (Fig. 3b). There is a risk of electrode (tungsten) migration to the workpiece if peak current greatly exceeds the average value. This technology exhibits moderate response. Newer technologies deliver a square wave output at line frequency (Fig. 3c). A square wave eliminates peaking and provides a rapid transition through zero, which is important to cyclic reignition of the arc. Thyristors are employed in concert with magnetic cores to generate the square current waveform. Adjustable imbalance permits the operator to control the ratio of electrode positive (EP) to electrode negative (EN) current by dwell extension (Fig. 3d). Direct current power sources can be classified as being constant-current sources or constant-voltage sources. Constant-Current Sources. For the SMAW process, a family of "drooping" volt-ampere (V-A) curves (A, B, and C) is

provided in Fig. 4(a). Often, a current boost or "dig" characteristic is made available (curve D in Fig. 4a). A burst of extra current during the brief period of metal transfer shorting tends to alleviate sticking during keyhole welding or when weaving or whipping the electrode, especially with the more-globular transfer electrodes.

FIG. 4 VOLT-AMPERE CURVES OF TYPICAL DC POWER SOURCES. (A) CONSTANT-CURRENT SOURCE. (B) CONSTANT-VOLTAGE SOURCE

Power sources for dc GTAW processes are generally more critically constant-current than are the SMAW "stick droopers." Ripple content is lower and current regulation is tighter. Inverter types customarily regulate to ±1% or ±1 A, whichever is larger. Remote foot control is essential to ensure that both hands are free for manipulation. Response time need not be fast, except for pulsed programs. Programmable and pulse-capable units permit the automation of complex current profiles. Power sources for GTAW processes include a high-frequency arc starting device that impresses a high radio frequency (RF) voltage on the electrode. This energy "jumps the gap" from the electrode to the workpiece, ionizing the shielding gas, and permits establishment of an arc. Thus, the electrode need not touch the workpiece. Constant-Voltage Sources. Power sources intended for gas-metal arc welding (GMAW) exhibit a relatively flat V-A curve, as shown in Fig. 4(b) (curves E, F, and G). The self-correcting characteristic of GMAW regulates the electrode burn-off rate. As the wire feed rate is increased, additional current will be drawn, but terminal voltage will remain

essentially constant. Voltage feedback techniques are used in solid-state designs to hold long-term output close to a set value. In the short-circuiting GMAW process, for example, drops of metal are transferred to the base metal through a pinch-off routine during a period of short circuiting of the electrode to the workpiece. Shorting time and arc time are determined by power source dynamics. Source Characteristics Whether utility fed or motor driven, all welding power sources have a characteristic that indicates their dynamic and steady-state response. In early machines, source characteristics were optimized for a single process and were relatively unchangeable. More recent designs can be programmed for a variety of characteristics that include multiple weld processes. The welding arc process is a dynamic, fast-changing load, exhibiting short circuits (if shorting, then no arc) at high currents and momentary voltages from near zero to over 50 V. Changes occur in milliseconds, particularly when metal transfer takes place. The ability of the power source to respond properly to rapidly changing conditions at the arc is critical to good performance. Modern technology invariably includes a closed-loop feedback circuit in the control system. Open-Circuit Voltage (OCV). When no load is connected to the output terminals of a welding power source, the

voltage that appears at the terminals is at its maximum. A high OCV value generally aids in arc starting and stability. In transformer-type power sources. OCV is established by the incoming utility line voltage and the transformer primary-tosecondary turns ratio. The output winding turns, the flux, and the angular velocity (rev/min) will establish the OCV in generators. In manual welding, the maximum OCV is 80 Vrms for machines with more than 10% ripple, or 100 V OCV for a ripple content of less than 10%. Power sources intended for automatic welding only are permitted 100 V OCV. The percentage ripple voltage is the ratio of rms ripple to average pulsating dc, expressed as: %ripple =

Vmax − Vmin Vr m s

(EQ 3)

Static V-A Curves. Two terms are commonly used to define the characteristics of a power source for welding: static

output and dynamic response. Static characteristics are set forth by the manufacturer as a series of V-A curves (Fig. 4). The slope of the static V-A curve is the ratio of load voltage change to the change in load current, expressed in V/100 A. It indicates the achievable envelope of output limits, including the highest voltage the machine will support at a given current, and the highest current that can be drawn at a given voltage. Dynamic Response. The static V-A curves discussed above do not take into account the dynamic interplay between the

power source and the process. Instantaneous response of the power source to step changes in load voltage is important to welding performance. This is especially true in the case of certain GMAW processes, where the response to a metal transfer short circuit is instrumental in bead formation and the reduction of spatter generation. Methods of Control or Electrical Conversation. Power source control methods are a key factor in the rate of current

rise (dI/dt). Either increased reactance in the ac circuitry or inductance in the dc portion will slow the rate of rise (and fall) of current during changes at the arc. Some equipment incorporates an adjustment to select the proper number of turns for a given welding condition (with silicon-controlled rectifiers [SCR] and transistor controls, these characteristics can be synthesized). Because there is no industry-accepted method to quantify the dynamics, the user must obtain a power source with a satisfactory range of adjustment. Another consideration is the frequency response of the power source, that is, its ability to "follow" a fast command signal, such as a pulse program. Magnetically based machines are generally not fast enough to faithfully reproduce a pulsed GMAW control signal and deliver it at high currents to the weld. When a better response is required, inverter-based SCR, transistorized sources with a high-frequency response are preferred. Droopers. A constant-current, a drooper, power source displays V-A curves such as those in Fig.(a) (curves A and C). A

considerable negative slope is evident. The internal impedance of the power source is said to be high. Large variations in

the arc (load) voltage (for example, below 40 V) do not appreciably affect the selected current. Droopers are noted for their use in the SMAW process, where the operator sets current to establish a suitable burn-off rate for the selected electrode. Manipulation, that is, shortening or lengthening the arc, does not greatly affect the arc current. However, if welding is conducted on a flatter V-A curve (see curve B in Fig. 3a), then electrode manipulation may change the current considerably. Thus, the operator gains more local control of the current through eye-hand coordination. Increasing the arc length by "pulling back" on the electrode causes the operating point to shift up slightly in terms of voltage, but decrease considerably in terms of current. Open-circuit voltage can be field adjusted in generator types, which provides the operator with a great degree of control through electrode manipulation. Skilled operators use this flexibility to their advantage when performing difficult out-ofposition welds. Constant-Voltage Sources. Equipment designed for constant-voltage processes will display a "stiff," nearly horizontal V-A curve (curve E, F, and G in Fig. 4b). The open-circuit voltage is 43 V for curve F. When serving a 300 A load, it falls to 35 V. This is expressed as: (43 − 35) 8V 2.6V = = 300 A 300 A 100 A

(EQ 4)

Curve H of Fig. 4(b) is representative of an inverter source in the constant-voltage mode. When set to 30 V, the electronic regulation system keeps the output very constant out to 300 A under a static (resistive) load, beyond which it falls abruptly to 0 V at 400 A. Curve K in Fig. 4(b) represents an extended output that may be available on a short-term (dynamic) basis from inverters. The term "constant" is taken to mean relatively constant. Some droop in voltage is permitted, and is even preferred in certain welding situations. Load current at a given voltage is responsive to the rate at which a consumable electrode is fed into the arc. The designer ensures that terminal voltage is maintained during heavy loads by providing tight coupling between the primary and secondary windings of the transformer, or by providing electronic regulation systems, or both. Movable Magnetic Cores. Welding power must be delivered to the arc under complete control of the operator. In the

simplest form, conversion to the lower voltage and selection of current are incorporated into a single, special welding transformer. Movable coils (varying the coupling between primary and secondary windings) and moving iron shunts provide stepless adjustment of the output (Fig. 2a). Saturable Reactors. In heavy industrial equipment, control is imparted by saturable reactors. These reactors comprise

laminated iron core assemblies fitted with power and control windings. A small control current varies the effective magnetic permeability of the iron core and controls flux to the power winding. Considerable gain is obtained. Typically, 5 A of control will command several hundred amperes of welding current. Both ac and dc outputs are available. Direct current output is obtained by the addition of secondary rectifiers. By combining rectifier diodes with self-saturating magnetic cores, a wider range is obtained, referred to as magamp control. Alternating current and ac/dc power sources are manufactured in sizes that range from 180 A Class III "farm welders" to units designed for industrial use and rated in hundreds of amps. Principal attributes are low cost and simplicity. Alternating current is specified wherever arc blow is a concern. It is also preferred for welding aluminum. Direct current machines are favored for most SMAW and GTAW processes, and for essentially all GMAW processes. Silicon-Controlled Rectifiers. In addition to the transformer taps, moving coils, moving shunts, and saturable reactors

mentioned earlier, there are a number of other means by which output can be controlled. Thyristors, commonly known a SCRs, are frequently employed as control elements, replacing magnetic cores. The SCR is similar to a rectifying diode, but is equipped with an additional gating element. With properly phased lowpower gate signals, a configuration of SCRs will deliver hundreds of amperes of current. The output of such a phase-controlled arrangement (Fig. 2b) can be tailored to perform as a stiff (that is, flat slope) constant-voltage source or as a drooping constant-current source through closed-loop feedback controls. When designed for three-phase operation, the dc output ripple frequency is either three times the line frequency (using three SCRs) or six

times, using six SCRs. Output ripple is "smoothed" with an appropriate inductor. Single-phase versions display the ripple of twice the line frequency and require a larger output inductor, L. Because welding current is analog in nature, it has traditionally been adjusted with analog circuitry via rheostats and potentiometers. Recent advances allow digital control with microprocessors and dedicated software; tight-tolerance repeatability is the driving force. Inverters. Inverter welding power sources are smaller and more efficient, and they offer better performance than their

line-frequency counterparts. The size and weight of iron-core components are reduced when operating at higher frequencies (Fig. 5).

FIG. 5 300 A RATED CONVENTIONAL LINE-FREQUENCY POWER SOURCE (RIGHT) AND INVERTER POWER SOURCE WITH SAME RATING (LEFT). COURTESY OF MILLER ELECTRIC MANUFACTURING COMPANY, INC.

With reference to the inverted shown in Fig. 2(c), the incoming utility line directly feeds a rectifier for conversion to direct current, which powers an inverter stage operating in the range from 2 to 100 kHz. Various topologies include forward, buck, series capacitor, and series resonant inverters. Control methods employ either pulse width modulation (PWM) or frequency modulation (FM). A variety of semiconductors are used as switching elements, including transistors (for example, field-effect transistors [FETs] and insulated-gate bipolar transistors [IGBTs]). The transformer, which is often designed of ferrite core material, operates at high frequencies without introducing high losses. The turns ratio determines welding voltage, usually 80 Vrms ac at the secondary winding. A second rectifier converts it back to dc for welding. A small output inductor is required to smooth the high-frequency ripple. Inverters have the greatest application where portability or advanced performance is required. Fast response permits current with fast transitions, low overshoot, and faithful reproduction of pulse programs. Variable Polarity. High-technology power sources permit adjustment of the ac waveform to vary the ratio of EP to EN

current and terms of both width and amplitude (Fig. 3e). Among these are switched dc types, wherein two dc constantcurrent power sources are alternately switched on and off, and arc current is of one polarity and then the other polarity. Almost unlimited frequencies can be generated, with rapid zero crossings. Cycloconverter power sources constitute a family that uses three-phase inputs to generate low-frequency ac for the

GTAW process. Through careful selection of specific portions of each line-frequency cycle, a pseudo square wave is assembled at a frequency lower than line frequency. Rotating Equipment. For a variety of reasons, rotating generators and alternators are often used to supply welding

power. In what is known as a motor-generator (M-G) set, an electric motor (Fig. 2d) drives the shaft of a generator, which also isolates weld power from the input utility power. The flywheel effect of the M-G set ensures stable output.

Where no electric utility power is readily available, engine-driven equipment can supply welding power (Fig. 2e). Gasoline and diesel fuels predominate. Both air- and water-cooled prime movers are available. Some include an auxiliary air compressor for air tool use. Generators produce direct current. A wire-wound armature spins in a frame arranged with electromagnetic field poles. The armature windings cut through these fields cyclically, generating ac. This is collected by commutator, that is, copper segments fitted to the armature shaft and equipped with fixed carbon brushes. The commutator serves as a "mechanical rectifier," converting the armature currents to dc. Both the OCV and current can be user adjusted, providing good control of the V-A slopes. The output is relatively pure dc, with very little ripple content. Engine speed is governor controlled. Most generators provide constant-current output for SMAW and GTAW processes, but adapters are often available for constant-voltage operation in the GMAW process. Auxiliary power for lights, tools, and so on derived from auxiliary stator windings. Alternators (that is, ac generators) produce alternating current, as the name implies (Fig. 2e). A spinning rotor winding is charged with direct current obtained from the exciter and applied through slip rings, creating a rotating magnetic field. No commutator is required. The rotating field cuts through a surrounding stator assembly, the windings of which collectively produce alternating current, either single phase or three phase, as designed. In single-phase designs, the resulting current can be used for ac welding. Output selection is provided in course ranges with reactor taps (Fig. 6), and fine adjustment is provided by rheostatcontrolled current in the revolving field windings. Engine speeds are generally either 1800 or 3600 rev/min, corresponding with four-people and two-pole alternators, respectively. Some alternators eliminate the revolving field winding by employing rotating permanent magnets. If direct current is required, a static (silicon diode) rectifier is included. The combination is referred to as a dc generator-rectifier.

FIG. 6 SCHEMATIC SHOWING TYPICAL ENGINE-DRIVEN WELDING POWER SOURCE

The recent evolution to solid-state controls has raised the performance level of engine-driven machines. Accessories include remote controls, idle devices to reduce engine speed when not welding, battery charging capabilities, and highfrequency arc igniters. Pulsed Power Supplies Pulsed GTAW and GMAW processes require a power source with the ability to deliver precisely shaped current pulses superimposed on a lower background (arc-sustaining) level. In addition to thyristors in phase control, inverters are also used to generate pulses and are especially versatile for such applications. Add-on "pulse formers" have also been designed for attachment for conventional dc welding power sources. Pulse Repetition Rates (PRR). For the GTAW process, PRRs (the number of times per second that a pulse is

transmitted) are relatively low (typically, 1 to 10 PRR). Fast rise times are not essential. However, in the GMAW pulse

process, repetition rates typically range from 100 to 200 PRR. Rise (and fall) times must be short (50 to 500 overshoot must be low to preserve pulse fidelity.

s) and

The mean current of a train of rectangular pulses is given by:

IAV = FTP(IP - IB) + IB

(EQ 5)

where Iav is the average current, IB is the background current, IP is pulse peak current, f is frequency, and TP is pulse width. From Eq. 5, with background current, pulse width, and peak current selected, the average current will be proportional to frequency. Increasing and decreasing pulse frequency in proportion to wire feed speed will balance the current with the burn-off rate of the electrode, maintaining the proper arc length for the GMAW process. Individual droplet size will remain essentially constant. Some pulse power sources generate pulses at line frequency or multiples thereof and stabilize the process by adjusting the power envelope of each pulse. (In a refined development, designated "one-knob control," the operator need only select a wire and gas type; pulse parameters are established by algorithms in the controller. As wire feed speed is adjusted, all other functions "track.") Synergic Pulsed GMAW. A step higher in technology results in synergic pulsed GMAW. Inverters (Fig. 2c) and

secondary switchers (Fig. 2f) ensure process stability by controlling melting rate through modulation of pulse shape and frequency. The relationship between digitally monitored wire feed speed and pulsation parameters is managed by microprocessor look-up tables and similar control strategies. The power source is simply required to exhibit fast response and high peak current capability. Multiple Operator Power Sources On large job sites with a number of arcs, a multiple operator (MO) power source can serve several operators. Diversity of use suggests that all arcs will not be operating simultaneously and that a 0.25 factor is typical. Accordingly, a single, large, 76 V constant potential power source can supply up to 20 SMAW operators. Each operator is equipped with a resistor "grid" to adjust arc current to his particular needs. An alternative MO concept provides individual magamp control modules (typically, eight) fed from a single, large, multiwinding transformer. Each operator can select his preferred polarity and current. Modules can be connected in parallel for greater output. Power Sources for Welding F. James Grist, Miller Electric Manufacturing Company, Inc.; William Farrell and Glen S. Lawrence, Ferranti Sciaky

Resistance Welding Power Sources The function of the power source for resistance welding (RW) is to deliver a predetermined amplitude of current to the welding electrodes that are clamping the workpiece. Current flows from the power source to one electrode, through the workpiece to the opposing electrode, and returns to the power source. The electrodes concentrate current into a small, usually circular area. The flow of this constricted current generates heat, concentrated largely at the interface(s) of the workpiece details to be joined. Heat Input. The rate of heating must be sufficiently intense to cause local electrode/workpiece interface melting. The

opposed electrodes apply a pressure and, when sufficient melting has been achieved, the current is interrupted. Electrode force is maintained while the molten metal solidifies, producing a sound, strong weld. If both the local resistance of the workpiece and the welding current magnitude were constant, then the total quantity of heat, Q, developed in the workpiece would be given (in joules) by:

Q = I2 RT

(EQ 6)

where I is the effective value of current in amperes, R is the resistance of the workpiece in ohms, and t is the duration of flow of current in seconds. However, the resistance is not constant. Contact resistances decrease in magnitude and the bulk resistance of the work column increases as its temperature rises. Furthermore, resistance variations of the workpiece may cause variations in the current amplitude, depending on the nature of the power source. Under these circumstances, the total quantity of heat, Q, is given by: t

Q = ∑ i ² r.dt

(EQ 7)

0

where i is the instantaneous value of current expressed as a function of time and r is the instantaneous resistance that is changing with time. Equipment Selection. The design of resistance welding machines resembles that of a "C" frame press (Fig. 7). The

power source transformer secondary winding is connected to the electrodes on cantilevered arms, one of which provides for adjustable electrode force. Throat depth is the distance from the frame to the electrodes. Throat gap is the vertical distance between the two parallel secondary conductors. The total area of the secondary circuit (throat depth times throat gap) exerts a profound influence on output performance. Any increase in inductance, whether due to geometry or magnetic materials within the secondary, will degrade performance and lower the power factor.

FIG. 7 SCHEMATIC SHOWING PRIMARY COMPONENTS OF A RESISTANCE WELDING MACHINE

The input to resistance welding power sources is usually 230 or 460 V utility power delivering single-phase or threephase alternating current at 60 Hz. All systems require transformers to step down the line voltage to a relatively low value, with a proportional increase in current. The simplest systems consist of a transformer connected to single-phase power. More advanced sources produce a dc welding current through rectification. Rectification Systems. In primary rectification systems, a special three-phase transformer is designed with a very large iron core. Each dc pulse is limited in duration and each is reversed in polarity to prevent saturation of the core. In the majority of applications, a weld is produced by a single impulse of direct current. For heavy gages, pulsation welding is used. A series of dc impulses of alternating polarity is separated by "cool time." The result is a very low frequency alternating current. Hence, the term frequency converter is commonly applied to this type of power source. Figure 8 shows a three-phase half-wave primary rectified power source.

FIG. 8 CIRCUIT DIAGRAM OF THREE-PLATE HALF-WAVE RESISTANCE WELDING POWER SOURCE WITH PRIMARY RECTIFICATION

Secondary rectification systems are available in single-phase full-wave, three-phase half-wave, and three-phase full-wave forms; the latter is most commonly utilized. Three-phase full-wave systems consist of three single-phase transformers, with rectifiers connected to the secondary windings (Fig. 9). With this system, current pulse time is limited only by thermal considerations and polarity is fixed.

FIG. 9 CIRCUIT DIAGRAM OF THREE-PHASE FULL-WAVE RESISTANCE WELDING POWER SOURCE WITH SECONDARY RECTIFICATION

The energy efficiency of a resistance welder depends on the efficiency of delivered energy at the electrodes. Energy is dissipated in the form of heat, because of resistance in the conductors. The current, I, which produces the heat of welding, is simply the secondary voltage divided by the total machine impedance, Z, plus that of the weld: i=

E Z

(EQ 8)

where Z is the vector sum of the resistance and the reactance: Z = R ² + X ² = R ² + (ω L)²

(EQ 9)

where R is the total resistance of the circuit and X = ωL is the total reactance of the circuit (where ωequals 2πf and L is the inductance). Resistance welders designed for single-phase utility power are lowest in investment cost, but highest in energy cost. The high-magnitude alternating current results in large inductive losses and a low power factor. Three-phase primary rectification and three-phase full-wave secondary rectification power sources reduce the power demand to about one half and operate at a higher power factor.

The rapidly pulsing weld zone temperature inherent with a single-phase source is least desirable from a welding standpoint. In addition, copper electrodes suffer from clamping impact, force, and heat, causing the occurrence of "mushrooming," which lowers current density. Mushrooming also reduces resistance in the current path. Welding current will increase, because current is controlled indirectly in resistance welding machines by adjusting secondary voltage (that is, advancing or retarding the firing angle). From Eq 8, a decrease in impedance causes an increase in the current. Therefore, all resistance in the work path influences the current magnitude produced by a given secondary voltage. This influence is small if total system impedance is high, but significant if the system impedance is low. A three-phase primary rectification system exhibits the greatest degree of this self-compensating effect, followed to a somewhat lesser degree by the three-phase secondary rectification category. Finally, a thermoelectric effect is associated with passing current from a copper alloy electrode to a workpiece of a given material. This phenomenon is important when welding aluminum or magnesium alloys with a single impulse. Reversing the current polarity for each successive weld, an inherent characteristic of the three-phase primary rectification system, provides the maximum electrode life between dressings. Duty Cycle. Resistance welding is inherently an intermittent process with a series of very short current periods followed

by "off" or "cool" periods. Therefore, this type of power source is rated on a 50% duty cycle and a 1 min integrating period. The relationship between kVA rating at 50% duty cycle and kVA demand at actual duty cycle is as follows: KVA50% = KVADemand 2.( D.C.)

(EQ 10)

where D.C. is the duty cycle in %, entered in the form of a decimal. Duty cycle is defined by: D.C. =

N .T 60. f

(EQ 11)

where N is the number of welds per minute, T is the weld time in cycles, and f is the frequency in hertz. Heat Controls. Resistance welds are short and are measured in cycles. Because timing is very important, digital

microprocessor controls abound. Electrochemical contactors have been replaced by SCRs or ignitrons connected in inverse parallel to handle the high currents. Coarse heat adjustment is accomplished by transformer tap switch setting, and fine adjustment is achieved through phase shifting, with digital circuitry providing the highest repeat accuracy. Additional information on resistance welding is available in the Section "Joining Processes" in this Volume. Power Sources for Welding F. James Grist, Miller Electric Manufacturing Company, Inc.; William Farrell and Glen S. Lawrence, Ferranti Sciaky

Electron-Beam Welding Power Sources Electron-beam welding (EBW) requires a variety of power supplies in two categories: those that generate and control beam position and those that control beam power. The outputs of various power supplies are coordinated by a machine control system. Beam Position and Intensity. Beam position arrangements for alignment, deflection, and focus are dominated by magnetic devices in electron-beam welders. Because the magnetic field effect on the beam (the magnetomotive force) is directly proportional to the current in the windings of the device, these power supplies are typically current regulated.

MMF = N · I

(EQ 12)

where N is the number of turns in the coil and I is the current in the coil. The desired current value is programmed by potentiometer, computer, or function generator.

A second factor regarding the effectiveness of magnetic devices is the velocity of the electrons in the beam, as determined by the accelerating voltage. At the voltages used in EBW (typically, 50 to 150 kV), relativistic effects must be considered. In situations where the beam potential is changing, but it is desired to have the effect of the magnetic devices be constant, then the effective magnetic excitation, k, must be constant: K² =

( N .I )² V*

(EQ 13)

where NI is the magnetomotive force and V* is the relativistic voltage. The relativistic voltage can be expressed in volts as:

V* = V + (0.9785 × 10-6)V2

(EQ 14)

Sensors monitor the beam potential generated by the high-voltage (HV) supply and correct the control signal to the magnetic device power supply. Thus, the action of the device can be held constant as beam potential changes. This correction can be achieved by using computer software or an analog circuit. Such correction is often called square root compensation, because of the relationship indicated in Eq 13. Alignment coils are used for slight corrections to the x and y positions of the beam as it enters the magnetic focusing coil. Power supplies for these devices are typically potentiometer programmed by the operator in an alignment procedure. These supplies are characterized by a high degree of long-term stability, with minimal need for dynamic response. Selected installations employ alternate means for achieving alignment. The magnetic focus coil acts very much as a glass focusing lens would for optics. It takes a divergent electron beam and converges it at a more distant point along the beam axis. The beam focus point from the magnetic lens is roughly inversely proportional to the square of the lens excitation: f ∝

V* ( N .I )²

(EQ 15)

The magnetic lens supply is usually controlled by computer output in order to coordinate beam focus with the weld program. Manual control (operator offset) and square root compensation are also usually provided. These supplies are of high stability and moderate dynamic response. The deflection coil moves the beam along the x and y axes to position the beam over the surface of the workpiece. Deflection rates can range from dc to high audio frequencies. Thus, the deflection supply is a high-quality currentregulated stereo amplifier. Amplifier input controls allow computer input for coordinated movements of the part and beam, function generator inputs for beam deflection patterns (rasters, and so on), and dc offsets to patterns. These are highly stable current-regulated coils drivers with wide dynamic response. High-Voltage Control. The electron beam is generated in an electron gun (resembling a diode or triode of cylindrical symmetry) that generates a paraxial flow of electrons. The power in an electron beam comes from the high-voltage supply that generates the large voltage difference between the anode and cathode of the gun. Electrons are emitted from the cathode and, accelerated by the voltage, fall to the anode. The beam passes through a central aperture in the anode and drifts through various alignment, focus, and deflection elements to strike the workpiece (Fig. 10). Both the workpiece and the anode are typically at ground potential, whereas the cathode is highly negative.

FIG. 10 SIX INDIVIDUAL POWER SUPPLIES THAT CONTROL POSITION AND INTENSITY OF THE BEAM IN EBW PROCESS

Control and regulation of the high voltage is usually done on the input to the high-voltage transformer/rectifier oil-filled tank. The input control installations range from motor-driven variable autotransformers, motor-alternators with field controls, SCR switching or phase control, to power MOSFET (metal-oxide semiconductor field-effect transistor) and IGBT devices in various switch mode regulators. In a three-phase input unit, a standard technique is to place the switching control in the primary neutral conductor of the HV transformer (Fig. 11). The secondary circuitry is three-phase or six-phase, full-wave rectified, and filtered.

FIG. 11 CIRCUIT DIAGRAM OF A HIGH-VOLTAGE CONTROL IN NEUTRAL IN A THREE-PHASE INPUT UNIT

In an alternate arrangement, line voltage is converted to dc, which powers an inverter to produce high-frequency singlephase input to the high-voltage transformer. Control can be provided either by regulating the level of the dc supply or by controlling the inverter pulse width or frequency. The HV transformer can be a conventional step-up unit or a voltage multiplier. These supplies can be characterized as voltage regulated and input controlled, with a moderate response rate. Thermionic Emission. The cathode heating unit provides the energy to raise the cathode to thermionic electron emission

temperatures. This is usually a dc source. For beam-heated cathodes, discharge-heated cathodes, and others, it may be more complicated, to a point of supplying power to a second miniature gun. In a space-charge limited electron beam, the cathode temperature must be held high enough that temperature-limited emission (Eq 16) is greater than space-charge limited flow (Eq 17):

IP = AT2 E −eφ / kt IP = K(VP + µ VG)3/2

(EQ 16) (EQ 17)

In Eq 16 and 17, Ip is beam current, Vp is accelerating voltage, A is Dushman's constant, Vg is bias voltage, and T is temperature. The bias supply provides voltage to the control electrode and regulates the magnitude of the beam current. Bias supplies can be input controlled with ground-level sensing of beam current on the return of the HV supply, or the beam can be sensed and regulated at high voltage using optical isolation (Fig. 12).

FIG. 12 CURRENT FEEDBACK AND BIAS CONTROL AT HIGH-VOLTAGE POTENTIAL

Ground-level sensing systems tend to use high-frequency inputs from linear amplifiers or switch mode inverters in order to achieve response on the order of a millisecond. Submillisecond response bias supplies are achieved by sensing and controlling the beam current at the high-voltage output. Because bias voltage is in the low kV range, vacuum tubes are used for output control. Additional information on electron-beam welding is available in the article "Electron-Beam Welding" in this Volume. Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Introduction OF ALL PHASE TRANSFORMATIONS, few have been more widely observed and studied than the transformation of a liquid to a solid (that is, solidification). The process of solidification is the same in all cases, whether it is the freezing of water on a windshield or in a freezer, or the solidification of metal in a casting or in the weld that joins two solids. The process is controlled by the free energy of the liquid phase, Gl, relative to that of the solid, Gs. This is depicted in Fig. 1, which shows the behavior of a pure (single component) material. Above the freezing temperature, Tf, the liquid phase has the lower free energy and is therefore stable, but below Tf, the solid is the stable phase. At Tf, both phases are in equilibrium, that is, Gl = Gs.

FIG. 1 TEMPERATURE DEPENDENCE OF BULK FREE ENERGY OF THE LIQUID AND SOLID PHASES IN SINGLECOMPONENT SYSTEM. THE SOLID LINE PORTION AND THE DASHED LINE PORTION OF EACH GL, AND GS CURVE INDICATE THE STABLE AND UNSTABLE PHASES, RESPECTIVELY, OF THE FREE ENERGY ON EITHER SIDE OF TF. SOURCE: REF 1

In the transition from one phase to another, the change in free energy, ∆G, is the difference in free energy of the product and the reactant. This free energy change can be expressed in terms of the enthalpy and the entropy changes, that is, for the transformation of a liquid to a solid during freezing:

∆G = GS - GL = (HS - HL) - T(SS - SL) = ∆H – T∆S

(EQ 1)

At the freezing temperature, Tf, ∆G = Gs - Gl = 0, because the free energy of the two phases is the same, and ∆H = Tf∆S, It is necessary to cool below Tf for solidification, because at Tf both the solid and liquid phases are present and in equilibrium. Below Tf, ∆G is not equal to zero (Fig. 1 shows that Gs < Gl) and is given by Eq 1 with T = T', where T' < Tf. Because the ∆H and ∆S are not strong functions of temperatures, they can be assumed to be temperature independent. Therefore, at any temperature, ∆H = ∆Hf and ∆S = ∆Sf, where ∆Hf and ∆Sf are the values of the enthalpy and the entropy changes for the equilibrium reactions at Tf (that is, the latent heat of fusion and the entropy change on fusion, respectively). Combining these enthalpy and entropy expressions, the fact that ∆H = Tf ∆S, and Eq 1, then at T', one obtains:

∆G = (∆HF/TF)(TF - T')

(EQ 2)

where ∆Hf, the latent heat of fusion, is negative. Hence, in agreement with Fig. 1, ∆G is negative. The greater the amount of undercooling (supercooling) below Tf (Tf - T'), then the greater the thermodynamic driving force for solidification. However, even when the conditions of Eq 2 are met, the liquid does not spontaneously transform to the solid below Tf. Rather, small amounts of solid nucleate and grow to produce complete solidification. Nucleation creates a new surface, that is, the surface between the solid and the liquid. The energy per unit are of this surface is the surface tension, γ, which is always positive. For solidification to occur, the increase in energy associated with the surface energy must be balanced by a greater decrease in the free energy of the solid relative to that of the liquid. This requires undercooling, as shown by Eq 2.

There are three ways in which a solid can form: homogeneous nucleation, heterogeneous nucleation, and epitaxial growth. Homogeneous nucleation occurs when there is no foreign body (mold wall, solid particle in the melt, etc.) on which to form the solid. Figure 2 shows the balance of the surface tension and the bulk free energy per unit volume, ∆Gv, as a function of the size of the nucleus that forms during homogeneous nucleation. The ∆Gv, value is just ∆G (as given by Eq 2) divided by the molar volume of the solid, Vs.

FIG. 2 FREE ENERGY OF FORMATION OF A NUCLEUS AS FUNCTION OF ITS RADIUS. SOURCE: REF 1

For a spherical nucleus of radius greater than rc, the volume free energy decrease outweighs the increase in energy that is due to the surface energy, and the nucleus is stable. At r = r*, the net energy is a maximum, but if additional atoms are added to the nucleus, then the energy decreases. The value r* is thus the radius at which the spherical nucleus is just stable, because as additional atoms are added, the energy is decreased. The value r* is given by:

R* = -2γ/∆GV

(EQ 3)

which is a positive number, because ∆Gv is negative. Substituting into Eq 2, one gets:

R* = -2 γTFVS/∆H F(TF - T')

(EQ 4)

The r* is positive because ∆Hf is negative and, for undercooling, (Tf - T') is positive. This equation is a general expression relating the radius of curvature of a surface and the degree of undercooling that is required for solidification of that surface. The greater the degree of undercooling, the smaller the radius of curvature that is stable. This equation will be considered again when nonplanar solidification is discussed. Nucleation requires a thermal energy either equal to or greater than ∆G* in order for a spherical nucleus of radius r* to form. The ∆G* value is the activation energy, which, as Fig. 2 shows, is a positive quantity. The greater the degree of undercooling, the more negative the free energy, the less positive the ∆G* value, and the greater the rate of nucleation of the solid. For homogeneous nucleation, ∆G* is given by:

∆G* = (16/3)(π γ 3TF2VS2)/(∆HF2∆T2) where ∆T = Tf - T'.

(EQ 5)

Heterogeneous nucleation develops when the solid forms on a foreign body. Now, the interaction of the nucleus and the foreign body must be considered. This interaction is defined in terms of the wetting angle between the nucleus and foreign body. The activation energy for heterogeneous nucleation is given by:

∆G* =

2 2 4π γ ³T f Vs (2 − 3cos θ + cos ³θ ) 3 ∆H 2f ∆T ²

(EQ 6)

The angle is 180° if there is no wetting, and Eq 6 reduces to Eq 5, that is, with no wetting, there can only be homogeneous nucleation. With a wetting angle of 90°, ∆G* is one half the value given by Eq 5. Instead of requiring a spherical nucleus of radius r*, a hemispherical nucleus of the same radius is necessary, which requires one half the number of atoms and one half the activation energy. This is the sort of nucleation that develops in a casting, where the mold wall acts as the foreign body. When the wetting angle is zero, Eq 6 shows that the activation energy is zero. This is the case for epitaxial growth on a substrate, where epitaxial derives from the Greek epi, upon, and taxis, to arrange, that is, to arrange upon. In effect, no new surface is being formed. Atoms are just being added to the substrate, thereby extending it. The important point is that no activation energy nor undercooling is required to add atoms onto an existing substrate. This is the situation that develops when a liquid solidifies on a substrate of the same material or one that is similar in composition and structure, as in the solidification of a weld.

Reference

1. J.D VERHOEVEN, FUNDAMENTALS OF PHYSICAL METALLURGY, JOHN WILEY & SONS, 1975

Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Comparison of Casting and Welding Solidification A weld can be thought of as a miniature casting. The fundamentals of weld solidification are the same as those of a casting, but with different boundary conditions. These differences and their effects are described below. First, by its very nature, a sound weld must attach to the metals being joined, whereas a casting must not adhere to the mold wall. To achieve this, the mold is treated to prevent adherence, whereas the joint to be welded is prepared to promote adherence. The practical result of this preparation is that there is generally excellent heat transfer through a weld joint, but relatively poor heat transfer through the mold, because oxides are often used as mold materials or for mold coatings (the molten metal will not wet the oxide and therefore cannot adhere to it). Second, heat is continually being added to the weld pool as it travels, whereas no heat is added to a casting after the pour, except for possibly modest heating of the mold. The practical result of this and of the first consideration described above is that the temperature of the casting is relatively uniform. In contrast, a very large temperature gradient develops in a weld. The center of a weld pool reaches a very high temperature (typically, 2000 to 2500 °C, or 3630 to 4530 °F), which is limited by the vaporization of the weld metal. At the sides of the weld pool, where solidification is taking place, the temperature is the solidification temperature, which is typically about 1000 °C (1800 °F) less than the temperature at the center on the weld pool. Thus, a positive temperature gradient is developed in a weld (that is, measured from the fusion line of the weld, the temperature increases going into the molten material). A casting will typically supercool to below the solidification temperature. Solidification will first develop by heterogeneous nucleation at the mold wall. The latent heat and the poor heat transfer through the mold wall will then cause a temperature rise at the mold wall. Thus, a small negative temperature gradient can develop briefly in a casting. The nature of this gradient is important, because it

promotes nonplanar solidification (to be discussed below) in the casting, whereas the positive gradient that develops in a weld helps to limit nonplanar solidification. Third, solidification is developed at the mold wall of a casting by heterogeneous nucleation, whereas epitaxial growth develops at the weld fusion line. Thus, some supercooling is required for the casting, but only a vanishingly small degree of supercooling is required for the weld. (As was discussed, epitaxial growth requires no supercooling, but some slight degree of undercooling is required to shift the reaction from an equilibrium between the solid and liquid, to all solid.) This undercooling in a weld would typically be less than 1 °C (1.8 °F). In contrast, in castings when there is a lack of effective nuclei for hetergeneous nucleation, undercoolings of 100 °C (180 °F) or more are possible. In general, however, the undercooling in castings is typically only on the order of a few degrees centigrade. Fourth, the generally larger volume of a casting, relative to that of a weld, and the poorer heat transfer makes the cooling rate and the solidification rate much lower for castings than for welds. Fifth, as a casting solidifies, the volume of the remaining liquid decreases. Thus, the shape of the molten pool is continually changing. In a weld, the weld pool shape is generally kept constant as it travels (if the heat input and section geometry are constant). Sixth, because of the stirring action of the arc and the action of Marongoni surface tension gradient induced convection forces, there is good mixing of the molten weld pool. In contrast, there is comparatively little mixing of the molten material of a casting. As discussed in the section "Solidification of Alloy Welds (Constitutional Supercooling)" in this article, all of these factors influence the nature of weld solidification, relative to what is observed in a casting. Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Solidification of Alloy Welds (Constitutional Supercooling) The solidification of an alloy is much more complex than that of a pure metal, whether in a weld pool or in a casting. Figure 3 shows a simple binary isomorphous phase diagram. The temperatures of the liquidus and solidus lines of Fig. 3 can be approximated by straight lines (for a limited temperature range). (The liquidus defines the temperature above which only liquid is present and the solidus the temperature below which only solid is present. The liquidus and solidus lines bound the two-phase liquid plus solid region.) The liquidus line is given by:

TL = TMA + MLCL

(EQ 7)

where TmA is the melting point of pure A, ML is the slope of the liquidus, and CL is the composition of the liquid at TL. The solidus is described by a similar expression.

FIG. 3 BINARY ISOMORPHOUS PHASE DIAGRAM SHOWING THE LIQUID AND SOLID COMPOSITIONS AT SELECTED TEMPERATURES

Another important quantity that can be estimated from the phase diagram is the distribution coefficient, K, which is defined as:

K = CS/CL

(EQ 8)

where CS and CL are the solidus and liquidus compositions, respectively, defined by equilibrium tie lines. Assuming that the solidus and liquidus lines are straight lines means that K is not a function of temperature. In general, the liquidus and solidus lines are curved, so that in actuality, K is a function of temperature. For simplicity, this curvature will be neglected and, instead, limited temperature ranges where K can correctly be assumed to be a constant will be considered. The use of a distribution coefficient defined by equilibrium tie lines implies that equilibrium exists between the liquid and the solid at the solid-liquid interface. This equilibrium need not exist elsewhere. Figure 3 and the following discussion are based on a K value of less than 1 and a negative ML value. The following discussion is also true for diagrams where K is greater than 1 and ML is positive. The distribution coefficient defines how the solute is distributed or partitioned between the solidifying solid and the remaining liquid. The slope of the liquidus line and the distribution coefficient are important in defining the composition gradient that develops when alloys solidify. In general, the distribution coefficient, K, is a function of temperature, but as already noted, to simplify the analysis, it will be assumed that over a limited temperature range, it is a constant. An equilibrium phase diagram, such as that of Fig. 3, assumes that equilibrium is maintained in all of the solid and liquid and that there are no composition gradients in the solid or liquid. It defines the equilibrium solid and liquid compositions at any temperature. Figure 3 shows that the solidification of an alloy of composition C0 will begin when the temperature is reduced below T1. If equilibrium is maintained, then solidification will be complete when the temperature is reduced below T3. A two-phase mixture of solid plus liquid will be present between T1 and T3, with the relative amount of each phase given by the lever rule:  C − Cs 2  %liquid =  0 100%  C1.2 − Cs 2 

(EQ 9)

at T2. The free energy curves corresponding to this equilibrium are much more complex than those of Fig. 1, which is relevant for a pure material.

The equilibrium predicted by Fig. 3 is never realized in practice, because it would require an infinitely slow cooling rate. When considering castings, it is generally assumed that the equilibrium predicted by the phase diagram is maintained at the solid-liquid interface, but that no composition change occurs in the solid as the temperature is lowered. At T1, a solid of composition CS1 forms. CS1 is equal to CL1K, and because at the liquidus CL1 = C0 the CS1 = C0K. When the temperature drops to T2, the solid that forms has a composition of CS2, and so on. A composition gradient is thus developed in the solid. As the solid forms, solute is rejected into the liquid, and it is assumed that this solute is mixed into the liquid, raising its composition to that predicted by the phase diagram. This occurs even though the mixing action in a casting is relatively poor, because the cooling rate is generally low enough to provide enough time to allow for the required solute redistribution by diffusion in the liquid. As a result, no gradient is developed in the liquid. The composition of the solid at the solid-liquid interface is still given by the equilibrium phase diagram. However, the previously formed solid is deficient in solute. As a consequence, complete solidification does not occur at T3. Further cooling is required, with solidification being completed only when the average solid composition is C0. The boundary conditions for a weld are somewhat different from those of a casting. The composition gradient that develops in the solid is assumed to be the same as that which forms in a casting, but the behavior of the liquid is assumed to be different. The mixing of the melt is greater than that developed in a casting, but the cooling rate is much higher. Thus, the solute that is rejected as the solid forms is not completely mixed into the liquid. Rather, the bulk liquid remains at the initial composition, C0, with a solute gradient formed in the liquid near the solid-liquid interface. At the liquids, T1, the solid that forms has a composition of CS1 = C0K. At T2, the solid composition is CS2. As with castings, it is assumed that the solid-liquid equilibrium predicted by the equilibrium phase diagram is maintained at the solid-liquid interface. Therefore, the liquid in equilibrium with this solid is CL2 = CS2/K. However, the lack of complete mixing means that the bulk liquid remains at C0, and that a composition gradient forms directly ahead of the solid-liquid interface. This composition gradient is illustrated in Fig. 4, which shows the composition gradient developed by unidirectional solidification and the temperature gradient driving this solidification. The solidification is modeled by assuming that a bar is being unidirectionally solidified, but this approach is applicable to solidification in general.

FIG. 4 INITIAL STAGE OF DIRECTIONAL SOLIDIFICATION OF A BAR, SHOWING THE TEMPERATURE PROFILE DRIVING THE SOLIDIFICATION AND THE COMPOSITION PROFILE ESTABLISHED FOR AN ALLOY WITH A SOLUTE CONTENT OF C0. SOURCE: REF 4

At temperature T3, the solid that forms has a composition of C0. The solid that is forming now has the same composition as the bulk liquid. Thus, an equilibrium situation exists. Solidification can now continue without having to further reduce the temperature. For castings, however, the composition gradient in the solid and complete mixing of the solute into the

liquid leaves the liquid still present at T3, thereby reducing the liquidus temperature. This means that, in a casting, the temperature must be lowered even more before further solidification can take place. Figure 4 shows the situation during weld solidification, whereas Fig. 5 shows the composition when solidification is complete. For most of the solidification, the solid forms at C0. The exceptions are at the start and at the end of solidification. The first solid that forms does so with a solute composition of CS1, which the phase diagram shows is less than C0. A composition gradient develops until temperature is lowered to T3, when the solid composition reaches C0 and an equilibrium is reached between the solid being formed and the composition of the bulk liquid. At the end of solidification, there is an increase in the solute content. This occurs to accommodate the solute buildup at the solid-liquid interface. At the end of the solidification, there is no longer any bulk liquid at C0. The solute that is ahead of the solidliquid interface raises the composition of the remaining liquid as it is rejected into this final small volume of liquid. This obviates the assumed boundary condition (a lack of mixing of the solute into the bulk liquid), and the situation becomes more like that of a casting (complete mixing in this small liquid volume). The final liquid composition of this final transient is the terminal composition predicted by the phase diagram, that is, the final liquid is driven to the composition of the lowest melting point. In the case of the simple binary isomorphous diagram used for this example, this is pure B. For a simple binary eutectic, it would be the eutectic composition.

FIG. 5 FINAL COMPOSITION PROFILES FOR THE UNIDIRECTIONAL SOLIDIFIED BAR SHOWN IN FIG. 4. THE INITIAL AND FINAL TRANSIENT ARE SHOWN FOR K < 1. R, RATE OF SOLIDIFICATION; DL, DIFFUSION RATE OF THE SOLUTE IN THE LIQUID. SOURCE: REF 4

The profile of the liquid composition gradient that is ahead of the solid-liquid interface is given by:

CL = CA e − [ RX L / DL ] + C0

(EQ 10)

where R is the rate of solidification. DL, is the diffusion rate of the solute in the liquid, Ca is the increase in the composition of the liquid at the solid-liquid interface (relative to C0), and XL is the distance into the liquid that is ahead of the solid-liquid interface. The liquid at the solid-liquid interface has a composition of CL3, which is in equilibrium with CS3 = C0. Therefore, CL3 = C0/K. The enrichment of the solute in the liquid, Ca, is thus (C0/K) - C0 or C0(1 - K)/K. The distribution coefficient, K, is thus an important variable in determining the composition profile in the liquid and, as will be shown, in determining the nature of the solidified microstructure. Substituting Eq 10 into Eq 7, one gets the expression describing the temperature variation of the liquidus that is ahead of the solid-liquid interface, namely:

TL = TMA + ML CA e − [ RX L / DL ] + ML C0

(EQ 11)

Figure 6 illustrates this liquidus profile, along with three possible temperature gradients that could actually exist in the liquid. They are all positive, reflecting the positive gradient that develops in a weld. Gradient A is greater than the slope of the liquidus at the solid-liquid interface, but those of B and C are not. The slope of the liquidus can be found by differentiating Eq 11 at XL = 0, that is, at the solid-liquid interface. For the temperature gradient to be less than this slope, the gradient G must be:

G < MLRC0(K - 1)/DLK

(EQ 12)

FIG. 6 LIQUIDUS PROFILE OF A SOLID-LIQUID INTERFACE. (A) LIQUIDUS PROFILE FOR STEADY-STATE CONSTITUTIONAL SUPERCOOLING. (B) THREE LIQUIDUS TEMPERATURE GRADIENTS (CURVES A, B, AND C) COMPARED TO THE LIQUIDUS PROFILE. SOURCE: REF 6

Equation 12 is important because it describes the temperature gradient required for constitutional supercooling (Ref 3, 4, 5). The establishment of the composition gradient in the liquid that is ahead of the solid-liquid interface produces a variation in the temperature of the liquidus. For curves B and C, this has allowed the actual temperature to be below the liquidus (that is, supercooling has developed, even though the temperature of the bulk liquid is above the bulk liquidus temperature). This is constitutional supercooling, and as will be shown, this promotes nonplanar solidification, even though there is a positive temperature gradient. The liquidus profile shown in Fig. 6 is correlated with the propensity for hot cracking. The greater the temperature difference between that of the solid-liquid interface and that of the bulk liquid, the greater the tendency for hot cracking. This is discussed in more detail in the article "Cracking Phenomena Associated with Welding" in this Volume. The constitutional supercooling model for describing weld solidification is presented because it qualitatively describes the evolution of different microstructures (described in the section "Development of Weld Microstructures" in this article). Many newer nonequilibrium approaches to solidification give better quantitative results. These approaches will be

discussed in the section "Nonequilibrium Effects: High-Rate Weld Solidification and Composition Banding" in this article.

References cited in this section

3. J.A. BURTON, R.C. PRIM, AND J. SLICHTER, J. CHEM. PHYS., VOL 21, 1953, P 1987-1991 4. W.A. TILLER, K.A. JACKSON, J.W. RUTTER, AND B. CHALMERS, ACTA METALL., VOL 1, 1953, P 428-437 5. W.A. TILLER AND J.W. RUTTER, CAN. J. PHYS., VOL 34, 1956, P 96-121 6. W.F. SAVAGE, E.F. NIPPES, AND T.W. MILLER, WELD. J., VOL 55, 1976, P 165S-173S Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Development of Weld Microstructures Nonplanar solidification develops when a protrusion moves ahead of the rest of the solid-liquid interface and continues to grow in a stable manner. This increases the surface area. The stable radius is given by Eq 4, which shows that the protrusion must move into a supercooled region in order to be stable. This is possible in a casting, because a negative temperature gradient is developed (that is, the temperature decreases from the solid-liquid interface into the liquid). In a pure material (single-component system), the positive temperature gradient that is established in the liquid of a weld prevents this nonplanar growth. In an alloy, however, constitutional supercooling allows nonplanar solidification in a weld, even with a positive temperature gradient. The lower the gradient, G, (compare curve C with curve B in Fig. 6), the greater the degree of constitutional supercooling. This is illustrated in Fig. 7, which shows several temperature gradients (Fig. 7f) compared to the variation in the liquidus and the nature of the solid-liquid interface (Ref 5, 6, 7). For a curve a, the temperature gradient is steeper than the liquidus temperature curve and only planar solidification is possible. For curve b, there is a shallower gradient and the solid-liquid interface can move by cellular growth. The gradient is successively lower for c and d, which allows larger protrusions and the development of cellular dendritic and columnar dendritic growth. At curve e, the gradient is so shallow and the resulting liquidus temperature is so far above the actual temperature in the liquid that equiaxed dendrites can form ahead of the solid-liquid interface.

FIG. 7 SCHEMATICS SHOWING MICROSTRUCTURE OF SOLID-LIQUID INTERFACE FOR DIFFERENT MODES OF SOLIDIFICATION AND THE TEMPERATURE GRADIENTS THAT GENERATE EACH OF DIFFERENT MODES. (A) PLANAR GROWTH. (B) CELLULAR GROWTH. (C) CELLULAR DENDRITIC GROWTH. (D) COLUMNAR DENDRITIC GROWTH. (E) EQUIAXED DENDRITE. (F) FIVE TEMPERATURE GRADIENTS VERSUS CONSTITUTIONAL SUPERCOOLING. SOURCE: REF 6

Figure 8, which is based on experimental observations, relates the gradient, solidification rate, and composition to the type of structure developed (Ref 5, 6, 7). The basis for this figure lies in Eq 12. If R is moved to the left side of Eq 12, then the right side contains only material variables. For a given alloy system (where DL, ML, and K are fixed), the greater the C0, the greater the tendency for constitutional super-cooling and nonplanar solidification. Equation 12 predicts that the

smaller the G/R, the more nonplanar the solidification. Figure 8 shows that, experimentally, the correlation is with G/ R , rather than G/R.

FIG. 8 PLOT OF SOLUTE CONTENT VERSUS SOLIDIFICATION PARAMETER TO SHOW THE NATURE OF SOLIDIFICATION. SOURCE: REF 6

As the dendrite, or columnar grain, grows into the liquid, it should follow the behavior shown in Fig. 5. (As will be discussed in the section "Nonequilibrium Effects: High-Rate Weld Solidification and Composition Banding" in this article, the experimentally determined compositions do not always exactly agree with that predicted by Fig. 5.) The center of each cell or dendrite behaves like the initial transient, that is, the solute content is low. The regions between the cells or dendrites behave like the final transients, that is, they are enriched in solute. This is illustrated schematically in Fig. 9. Cellular growth occurs for a smaller C0 and/or for a larger G/ R than that for dendritic growth (see Fig. 8).

FIG. 9 SCHEMATIC SHOWING SOLUTE DISTRIBUTION AT THE DENDRITE OR CELL CORE AND IN THE INTERCELLULAR OR INTERDENDRITIC REGIONS. (A) CELLULAR GROWTH. (B) DENDRITIC GROWTH. SOURCE: REF 8

The constitutional cooling described by Fig. 4, 5, 6, 7, 8, and 9 occurs on a very fine scale. Solving Eq 10 shows that the solute composition that is ahead of the solid-liquid interface will decrease to almost C0 when XL = 5DL/R. For a typical liquid diffusion rate of 5 × 10-5 cm2/s (5 × 10-8 ft2/s) and a solidification rate of 5 mm/s (0.2 in./s), the liquid composition drops to C0 within about 5 μm (200 μin.). In the initial transient, the distance over which the solid composition increases to C0 is given by 5DL/RK. Thus, for K < 1, this distance is greater than the region in the liquid just ahead of the solidliquid interface. This is illustrated in Fig. 5. This difference is also important in defining the extent of the final transient, which is of the order of the length of the zone in the liquid where the composition drops to C0. The length of the initial transient is thus 1/K larger than that of the final transient. The length of the solute-poor cell or dendrite core is thus 1/K larger than that of the solute-rich intercellular or interdendritic regions. The solute-rich region between the cells or dendrites has a lower freezing temperature than that of the cellular or dendritic core. Thus, it is possible to have liquid films between impinging cells or dendrites. These films cannot support the thermally induced shrinkage stress that develops during solidification, and hot cracking can develop (see the article "Cracking Phenomena Associated with Welding" in this Volume). The hot cracking produced by nonplanar solidification can be minimized by reducing C0 (that is, by using purer materials) and/or by increasing G/ R (see Fig. 8). The temperature gradient in the liquid, G, is inversely proportional to the heat input, Q, which is given by:

Q = FEI/V

(EQ 13)

where f is the efficiency of the welding process, E is the arc voltage, I is the arc current, and V is the welding torch velocity. Thus, G can be increased by decreasing either the arc voltage or arc current, or by increasing the torch velocity. The influence of the velocity is counterbalanced by the fact that the solidification rate, R, is proportional to V (see the section "Effect of Welding Rate on Weld Pool Shape and Microstructure" in this article). Thus, increasing V increases R, which tends to increase the degree of constitutional supercooling.

References cited in this section

5. W.A. TILLER AND J.W. RUTTER, CAN. J. PHYS., VOL 34, 1956, P 96-121 6. W.F. SAVAGE, E.F. NIPPES, AND T.W. MILLER, WELD. J., VOL 55, 1976, P 165S-173S 7. F. MATSUDA, T. HASHIMOTO, AND T. SENDA, TRANS. NATL. RES. INST. MET. (JPN), VOL 11, 1969, P 43 8. W.F. SAVAGE, C.D. LUNDIN, AND A.H. ARONSON, WELD. J., VOL 44, 1965, P 420S-425S Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Effect of Welding Rate on Weld Pool Shape and Microstructure The velocity of the welding torch affects not only the rate of solidification, but the shape of the weld pool and the propensity to develop centerline hot cracks. The shape of the weld pool is dictated by the velocity, V, at which the welding torch moves and by the rate at which heat can be removed at the solid-liquid interface. To keep a constant shape, the rate of new melting must be exactly balanced by the solidification rate. On the average, this solidification occurs normal to the solid-liquid interface, because this is the direction of the maximum temperature gradient and, thus, the direction of the maximum heat removal. Simple geometry requires that for the weld pool shape to remain constant, the solidification rate, R, must be related to the torch velocity by:

R = V COS φ

(EQ 14)

where φ is the angle between the normal to the solid-liquid interface and the direction of motion of the torch (Fig. 10). At the back of the weld pool, where φ = 0, R = V. If this were not so, then the weld pool would be elongating or shortening. At the side of the weld pool, where φ = 90°, R = 0. If this were not the case, then the weld pool would be decreasing in diameter. With a constant weld pool shape, solidification at the side of the weld pool occurs as the region adjacent to where φ = 90° sweeps by, rather than by the motion normal to the weld pool velocity.

FIG. 10 SCHEMATIC OF A MOVING WELD POOL SHOWING THE RELATIONSHIP BETWEEN VELOCITY OF TRAVEL OF WELDING TORCH, V, AND THE RATE OF SOLIDIFICATION, R, AT SELECTED POINTS ALONG WELD POOL BOUNDARY. SOURCE: REF 9

Equation 14 predicts that R will vary around the weld pool. Figure 8 shows that as R varies, so does the microstructure. The result of these two effects is a variation of the microstructure around the weld pool (Ref 7). Planar solidification is to be expected at the sides of the weld pool, with the extent of the nonplanar solidification increasing towards the center of the weld pool. This is shown schematically in Fig. 11.

FIG. 11 SCHEMATIC SHOWING VARIATION OF MICROSTRUCTURES IN RESPONSE TO VARIATION OF THE SOLIDIFICATION RATE AROUND THE WELD POOL. SOURCE: REF 7

The coarseness of the nonplanar structure is also a function of R. The dendrite tip radius varies inversely with the square root of R. Not only does the dendrite tip radius become smaller as R increases, the dendrite arm spacing does also, although the exact form of the dependency is more complex. The same general inverse dependency with R is also true for the cell spacing when the solidification is cellular. The ability of the solid-liquid interface to move at a velocity, V, depends on two factors: the maximum solidification rate, which is a function of the crystallographic orientation of the solid at the point in question, and the ability of the temperature gradient in the solid to remove the heat and establish the appropriate temperature for solidification to proceed. Consider first the influence of the orientation of the solid grain at the solid-liquid interface. For body-centered

cubic (bcc) and face-centered cubic (fcc) metals, the preferred growth direction for epitaxial solidification is the direction. For hexagonal close-packed (hcp) metals, the preferred growth direction is . The planar growth of grains oriented with this direction normal to the solid-liquid interface is favored. This is illustrated in Fig. 12. Grains with the preferred orientation will grow at the expense of adjacent grains that are less favorably oriented (Ref 8, 9, 10, 11, 12). Because dendrites follow the same preferred growth directions, those with the proper orientation will be favored. The requirement for a constant weld pool size becomes somewhat more complex when the effect of crystallographic orientation is considered (Fig. 12). Equation 14 can be modified to become:

R' = V COS φ /COS ( φ ' - φ )

(EQ 15)

where R' is the growth rate in the preferred direction required to maintain a constant weld pool shape and φ ' is the angle between R' and V, and ( φ ' - φ ) is the angle between the local preferred growth direction and R, which is normal to the solid-liquid interface.

FIG. 12 SCHEMATIC SHOWING MOVEMENT OF A CURVED SOLID-LIQUID INTERFACE FOR SEVERAL GRAINS (A AND B), AND THE CHANGE IN THE RAPID GROWTH DIRECTION RELATIVE TO THE INTERFACE POSITION.

Figure 12 shows the situation at successive positions of the weld pool as it moves in the direction of V. The situation for the weld moving from position 1 to position 2 is shown. The weld pool is assumed to be elliptical in shape, so that at any point (except at the very back of the weld pool where φ = 0) as the pool moves, the angle between R (or R') and V changes. If an adjacent grain becomes oriented such that R' becomes normal to the solid-liquid interface [that is, so that ( φ ' - φ ) is equal to zero], it will become most favored and will grow at the expense of adjacent grains. In Fig. 12, between positions 0 and 1, grain A is more favorably oriented than grain B. The reverse is true between positions 1 and 2. Figure 12 uses only a few widely spaced positions, but in reality there is a continuum of such positions.

Competitive growth causes grains to try to rotate into the direction of V. Figure 13(a) shows how the grains rotate in an elliptical weld pool. Dendrites will also try to rotate into the R direction, although the grains or dendrites do not actually rotate. Rather, the growth shifts to the most favored (or ) direction of the six present in the fcc, bcc, or hcp lattices. As grains are edged out by more favorably oriented grains, they get smaller and eventually can disappear. The solute that they were rejecting is spread out as this competitive process favors and disfavors different grains. Those grains or dendrites that eventually grow into the V direction will reject their solute into the weld pool, which gradually becomes solute enriched, contributing to possible crater cracking when the weld terminates. To combat this, the weld power should be gradually decreased in order to maintain a steep temperature gradient (minimizing the degree of constitutional supercooling). The use of a runoff weld tab could also be considered.

FIG. 13 SCHEMATIC SHOWING EFFECT OF HEAT INPUT AND WELDING SPEED VARIATIONS ON WELD GRAIN STRUCTURE. (A) LOW HEAT INPUT AND LOW WELDING SPEED, PRODUCING AN ELLIPTICAL WELD POOL. (B) HIGH HEAT INPUT AND HIGH WELDING SPEED, PRODUCING A TEAR-DROP-SHAPED WELD POOL. HERE, THE HEAT INPUT AND WELDING SPEED ARE NOT YET SUFFICIENT TO CAUSE HETEROGENEOUS NUCLEATION AT THE WELD POOL CENTERLINE. (C) HIGH HEAT INPUT AND HIGH WELDING SPEED, WITH HETEROGENEOUS NUCLEATION AT THE WELD POOL CENTERLINE. SOURCE: REF 13

The second factor to consider is the ability to extract heat at the solid-liquid interface. Heat is extracted by the temperature gradient of the solid. This gradient is a minimum at the back of the weld pool, where the temperature of the solid is the highest. Unfortunately, this is exactly where the solidification rate must be a maximum to keep up with V. The temperature gradient is a maximum at the side of the weld pool, where the least heat extraction is required, because here R = 0. As long as the maximum possible solidification rate, Rmax, is equal to or greate than V, the weld pool can maintain an elliptical shape. If V is increased above Rmax, then the solidification rate at the back of the weld pool cannot keep up, and

the weld pool elongates. Increasing V decreases the heat input, which shrinks the weld pool volume. With the weld elongating, this means that the diameter must decrease as the pool elongates. This causes the weld to take on a tear-drop shape, as shown in Fig. 13(b). The angle at the back of the weld pool will decrease until φ critical is reached. φ critical is given by: φ CRITICAL

= COS-1 (RMAX/V)

(EQ 16)

Rmax is controlled by the thermal conductivity of the solid, the geometry of the part being welded (primarily the thickness), and the preferred direction growth rate. When there is a shift to a tear-drop weld pool, the grain (or dendrite) rotation is limited to φ critical. When this situation is attained (Fig. 13b), the grains (or dendrites) will all grow together at the center of the weld pool. The solute that is being rejected into the terminal transient will build up at the centerline, which can lead to centerline cracking. This cracking is caused by the lower solidification temperature of this solute-rick centerline region. As the surrounding material solidifies, stresses are developed that cannot be supported by the liquid centerline material, causing centerline cracking. The obvious remedy is to reduce the welding rate, although this has a production rate penalty. Figure 13(b) typifies what is to be expected from a pure material or one with a low solute content. In the case of alloys or in cases when the solute content is high, Fig. 8 predicts that when R is increased, not only does the weld pool become tear-drop shaped, but the nature of the structure changes, as shown in Fig. 11. If R is large enough, equiaxed grains can develop from heterogenous nucleation in the supercooled centerline material (Fig. 13c). Such behavior has indeed been observed in aluminum alloys (Ref 13, 14, 15). Figure 14 shows actual weld microstructures corresponding to the schematic structures of Fig. 13.

FIG. 14 ACTUAL WELD MICROSTRUCTURES CORRESPONDING TO SCHEMATICS OF FIG. 13. (A) GRAIN STRUCTURE OF A GTAW OF ALUMINUM ALLOY 6061, MADE WITH Q OF 700 W (200 BTU/H) AND V OF 5.1 MM/S (0.20 IN./S). (B) GRAIN STRUCTURE OF A GTAW OF PURE ALUMINUM, MADE WITH V OF 20.8 MM/S (0.819 IN./S). (C) GRAIN STRUCTURE OF A GTAW OF ALUMINUM ALLOY 6061, MADE WITH A HIGHER Q THAN THAT OF (A) (1320 W, OR 385 BTU/H) AND V OF 12.7 MM/S (0.500 IN./S). GTAW, GAS-TUNGSTEN ARC WELD. SOURCE: REF 13, 15

References cited in this section

7. F. MATSUDA, T. HASHIMOTO, AND T. SENDA, TRANS. NATL. RES. INST. MET. (JPN), VOL 11, 1969, P 43 8. W.F. SAVAGE, C.D. LUNDIN, AND A.H. ARONSON, WELD. J., VOL 44, 1965, P 420S-425S

9. W.F. SAVAGE, WELD. WORLD, VOL 18, 1980, P 89-113 10. W.F. SAVAGE, C.D. LUNDIN, AND R.J. HRUBEC, WELD. J., VOL 47, 1968, P 420S-425S 11. W.F. SAVAGE, C.D. LUNDIN, AND T.F. CHASE, WELD. J., VOL 47, 1968, P 522S-526S 12. W.F. SAVAGE, E.F. NIPPES, AND J.S. ERICKSON, WELD. J., VOL 55, 1976, P 213S-221S 13. S. KOU AND L. LE, METALL. TRANS. A, VOL 19, 1988, P 1075-1082 14. T. GANAHA, B.P. PEARCE, AND H.W. KERR, METALL. TRANS. A, VOL 11, 1980, P 1351-1359 15. H. NAKAGAWA, M. KATOH, F. MATSUDA, AND T. SENDA, TRANS. JPN. WELD. SOC., VOL 4, 1973, P 11 Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Nonequilibrium Effects: High-Rate Weld Solidification and Composition Banding As has been noted, while the constitutional supercooling model (described in previous sections of this article) gives valuable insights into the establishment of nonplanar structures, it does not necessarily give an exact depiction of the composition gradient develop within the dendrites or cells that may form and within the interdendritic or intercellular regions. This has been shown by Brooks and Baskes (Ref 16) in their study of gas-tungsten arc and electron-beam aluminum-copper and iron-niobium welds. For the aluminum-copper welds, the cell core compositions were about 2 to 3 times the KC0 value expected for the first material to solidify (see Fig. 5) and almost 4 times KC0 for the iron-niobium welds. This discrepancy was ascribed to the undercooling because of the curvature at the cell tip (see Eq 4), which is not considered in the constitutional supercooling model. Tip undercooling decreases the temperature at which the solidification occurs and, as Fig. 3 shows, increases the composition of the solid that is forming. No initial transient of the sort shown in Fig. 5 was observed. This was explained by using the lower solidification rate at the side of a cell, rather than that at the dendrite tip in the calculation of the size of the initial transient. This increases the length of the transient (which is inversely proportional to the solidification rate, R) and eliminates the sharp composition variation of the initial transient shown in Fig. 5. Brooks and Baskes also incorporated postsolidification solid-state diffusion of solute (following the approach of Brody and Flemings, which is given in Ref 17) to explain the composition gradients that they measured. Such diffusion is also not considered in the conventional constitutional supercooling model. In addition, the classical theory of constitutional supercooling, as described by Eq 10, 11, and 12, does not take into account the temperature gradient that develops in the solid, the thermal conductivity of the solid and the liquid, and, most importantly, the surface tension of the solid-liquid interface. These factors are considered in interface stability models (Ref 18, 19, 20, 21, 22, 23, 24, 25, 26), which describe the breakdown of a planar solid-liquid interface in terms of the stability of nonplanar perturbations. The importance of these models is that they explain the observed transition from nonplanar solidification back to planar solidification when the solidification rate is increased to very high levels. As has already been shown, at low solidification rates there can be a transition from planar to nonplanar solidification, as described by constitutional supercooling (as shown in Fig. 8). As the solidification rate is increased, the radius of curvature and spacing of the dendrite arms will decrease, as does the cell spacing. What eventually occurs, at very high solidification rates, is that the very small radius that is required is not supportable, because of the increased surface energy. The stability models predict that there will be a solidification rate (the limiting stability rate) beyond which there will be a transition back to planar solidification. Figure 15 shows the calculated dendrite tip radius as a function of velocity of the solid-liquid interface. The increase in R at high rates defines the limiting stability rate.

FIG. 15 PLOT OF DENDRITE TIP RADIUS, R, VERSUS VELOCITY, V, OF THE SOLID-LIQUID INTERFACE FOR AN AG-5% CU COMPOSITION WITH TEMPERATURE GRADIENT, G, OF 105 K/CM (5 × 105 °F/IN). CURVE A HAS DL(T) AND K(V) WITH KE AND ML AS CONSTANTS; CURVE B, DL (T) WITH KE AND ML AS CONSTANTS; CURVE C, WITH DL, KE, AND ML AS CONSTANTS. THE VELOCITY WHERE THE TIP RADIUS INCREASES IS THE STABILITY LIMIT. THE CALCULATIONS ARE MADE USING THE CONVENTIONAL STABILITY LIMIT APPROACH, WITH THE ADDITION OF A TEMPERATURE-DEPENDENT DIFFUSION COEFFICIENT, AND WITH THE ADDITION OF BOTH A TEMPERATURE-DEPENDENT DIFFUSION COEFFICIENT AND A VELOCITY-DEPENDENT PARTITION COEFFICIENT. SOURCE: REF 21

Each of the three curves in Fig. 15 illustrates different assumptions. The right-most curve shows the result using the approach of Mullins and Sekerka (Ref 18), who developed the original interface stability model. Here, the solute diffusion rate in the liquid, DL, the equilibrium distribution coefficient, K, and the liquidus slope, ML, are all assumed to be constants. The middle curve takes into account the fact that the diffusion coefficient is temperature dependent. (Increasing the degree of undercooling that is due to rapid solidification decreases both the solute diffusion rate and the limiting velocity.) The left-most curve not only considers the temperature dependence of the diffusion coefficient, but also takes into account a dependency of the distribution coefficient on the solidification rate. The preceding discussions have utilized the distribution coefficient, K, determined from equilibrium phase diagrams. At high solidification rates, this equilibrium is not achieved at the solid-liquid interface. Rather, K is a function of the solidification rate. One of several different expressions for the variation of K with V is given by (Ref 23):

K = [KE + (V A0/DL)] / [1 + (VA0/DL)]

(EQ 17)

where Ke is the equilibrium distribution coefficient (K in Eq 8 and 12), and a0 is a distance that is related to the interatomic spacing. Different models use different definitions for a0 and different functional relationships, but give similar results. When V > D/a0, K will approach 1. For D = 5 × 10-5 cm2/s, or 5 × 10-8 ft2/s (a typical value for liquid diffusion), and a0 = 5 × 10-8 cm, or 2.0 × 10-8 in., K will approach 1 for V > 103 cm/s, or 4 × 102 in./s. With K = 1, there is no partition of the solute, no constitutional supercooling, and the solidification will be planar for a positive temperature gradient. Figure 15 shows that the limiting velocity, as predicted by the stability model, is quite a bit lower than 103 cm/s (4 × 102 in./s). Thus, the reestablishment of planar solidification is not just the result of K approaching unity. Rather, it is due to the lack of the stability of nonplanar protrusions, which is influenced by this change in K. This development of planar solidification is important because it can arise in electron-beam and laser welds, and because the development of planar solidification lessens the tendency for hot cracking. Figure 15 shows the results of calculations, whereas Fig. 16 shows the results of actual observations made on silvercopper electron-beam scans (passes of the electron beam). These results are in general agreement with the predictions of Fig. 15. Figure 16 shows the compositional dependence of the transition to segregation-free planar solidification. It is

analogous to Fig. 8, which describes the transition from planar to nonplanar solidification that occurs at lower solidification rates. Figure 16 also shows that just below the limiting solidification rate, a banded structure can develop. This banding develops at a constant solidification rate and is due to instabilities at the solid-liquid interface.

FIG. 16 OBSERVED MICROSTRUCTURES IN SILVER-COPPER ALLOYS AS A FUNCTION OF THE SCAN VELOCITY (BEAM TRANSLATION VELOCITY) FOR ELECTRON-BEAM WELDS. OPEN CIRCLES DENOTE MICROSEGREGATIONFREE STRUCTURE; CLOSED CIRCLES, A STRUCTURE CONSISTING OF CELLS OR DENDRITES; AND OPEN TRIANGLES, THE DEVELOPMENT OF A BANDED STRUCTURE. SOURCE: REF 24

A more common type of banding is associated with the ripples observed on the surface of most welds. Variations in the welding power will cause variations in the solidification rate. If the power decreases, then the solidification rate will increase. More solute will be incorporated into the solid and a solute-rich band will be created (for K < 1 and where ML is negative). The reverse is true for an increase in the power (that is, the solidification rate will decrease), and this will cause a solute-poor band to form (for K < 1 and where ML is negative). These power fluctuations can be deliberate, as in pulsed welding, or can result from unavoidable instabilities in the welding power supplies or fluctuations in primary voltage, mechanical (travel) motion, and so on. Solute banding and surface rippling are illustrated in Fig. 17.

FIG. 17 SOLUTE BANDING AND SURFACE RIPPLING IN WELD BEADS. (A) SURFACE RIPPLES IN A TUNGSTENINERT GAS WELD MADE WITH AL-5SI FILLER. 6×. (B) SULFUR AND PHOSPHORUS SOLUTE BANDING AT THE FUSION BOUNDARY OF A MANUAL ARC WELD IN CARBON-MANGANESE STEEL. 26×. COURTESY OF THE WELDING INSTITUTE

Variations in the welding speed can also cause banding, as can the waves associated with the weld pool circulation, which can also influence the solidification rate. Special alternating magnetic fields have been used to supplement the surface tension forces, the usual electromagnetic forces, and the arc pressure and buoyancy forces, all of which affect the weld pool circulation (Ref 27). The aim is to induce interface fragmentation and cause grain refinement. The solidified structure can also be influenced by several other factors, in addition to variations in the rate of movement of the solid-liquid interface (Ref 27). Inoculants have been used to nucleate equiaxed grains by heterogeneous nucleation on the inoculant particles. Ferro-titanium, ferro-niobium, titanium carbide, and titanium nitride have been used as grain refiners in steels. Zirconium and titanium have also been used as inoculants for aluminum. The effectiveness of inoculants depends not only on the type of inoculants used, but on the size of the inoculant particles, the method of introduction (where in the weld pool it is added), and welding conditions. Arc motion and vibrations have also been used to suppress columnar growth and to produce a finer grain size (Ref 27). The effectiveness of arc vibration on the as-solidified grain size of an Al-2.5% Mg weld is illustrated in Fig. 18. The effectiveness of this approach depends on the frequency and amplitude of the arc motion, the heat input, arc length, and sheet thickness. In general, the greater the amplitude of vibration, the less the frequency required for grain refinement (Ref 28).

FIG. 18 PLOT OF GRAIN SIZE VERSUS ARC VIBRATION AMPLITUDE FOR AL-2.5MG WELD. WELD PARAMETERS: WELDING CURRENT, 150 A; WELDING SPEED, 200 MM/MIN (50 IN./MIN); ARC GAP, 2.4 MM (0.094 IN.); ARC VIBRATION FREQUENCY, 10 HZ. SOURCE: REF 27

Campbell (Ref 28) has developed frequency-amplitude maps that define several physical effects associated with the use of vibrations (that is, peak pressures generated, development of shearing stresses, power densities, laminar flow around dendrite arms, and the development of standing waves). These factors affect microstructural features, such as grain size, amount of microsegregation, and the development of porosity. For a given frequency of vibration, increasing the amplitude of vibration can cause a refinement of the grain size (Fig. 18), reduce microsegregation, and, unfortunately, increase the degree of porosity that is caused by vibration-induced cavitation. Vibrations can also cause splatter of the liquid weld metal; coarsen intragranular features, such as the secondary dendrite arm spacing; and promote hot tearing. Great care must therefore be taken when using vibratory motion and when determining the optimum amplitude and frequency of vibration. Variations in the arc current can also help to refine the weld microstructure. Ultrasonic vibration of the weld pool has also been found to be effective in some cases. These vibrations have been added through the filler wire and through the use of water-cooled copper probes (in large weld pools, such as those developed in electro-slag welding). In some cases, variations in the nature of the weld pool circulation have also been found to be effective in refining the weld structure. The desire to produce a fine-grained equiaxed weld structure is driven by the improved mechanical properties and lessened hot cracking that this structure provides. Torch weaving (Ref 29) (that is, moving the arc in a sinusoidal rather than linear fashion) has been used to break up the linear nature of the weld and to prevent hot cracks from propagating along the weld. A weld made with torch weaving is shown in Fig. 19. This weaved weld structure prevents cracks from running the length of the weld, but adds to the cost of the weld by decreasing the linear travel speed.

FIG. 19 TYPICAL GRAIN STRUCTURE OBTAINED WHEN TORCH WEAVING IS USED TO WELD ALUMINUM ALLOY 2014. WELD PARAMETERS: TRANSVERSE ARC FREQUENCY OSCILLATION, 1.0 HZ; AMPLITUDE OF OSCILLATION, 1.5 MM (0.059 IN.). SOURCE: REF 13, 30

Our understanding of weld solidification owes much to the more extensive study of the solidification of castings. The reader should refer to that subject area for a more in-depth treatment of solidification in general. However, it should be remembered that the boundary conditions for weld solidification and for that of a casting are not the same. Therefore, care must be taken in applying all that has been learned from casting metallurgy to the metallurgy of welds.

References cited in this section

13. S. KOU AND L. LE, METALL. TRANS. A, VOL 19, 1988, P 1075-1082 16. J.A. BROOKS AND M.I. BASKES, WELD MICROSTRUCTURE CHARACTERIZATION AND MODELING, RECENT TRENDS IN WELDING SCIENCE AND TECHNOLOGY, S.A. DAVID AND J.M. VITEK, ED., ASM INTERNATIONAL, 1990, P 93-99 17. H.D. BRODY AND M.C. FLEMINGS, TRANS. AIME, VOL 236, 1966, P 615-624 18. W.W. MULLINS AND R.F. SEKERKA, J. APPL. PHYS., VOL 35, 1964, P 444-451 19. J.S. LANGER AND H. MULLER-KRUMBHAAR, ACTA METALL., VOL 26, 1978, P 1681-1687 20. W. KURZ AND D.J. FISCHER, ACTA METALL., VOL 29, 1981, P 11-20 21. W. KURZ, B. GIOVANOLA, AND R. TRIVEDI, ACTA METALL., VOL 34, 1986, P 823-830 22. R. TRIVEDI AND W. KURZ, ACTA METALL., VOL 34, 1986, P 1663-1670 23. M.J. AZIZ, J. APPL. PHYS., VOL 53, 1982, P 1158-1168 24. W.J. BOETTINGER, D. SHECHTMAN, R.J. SCHAEFFER, AND BIANCANIELLO, METALL. TRANS. A, VOL 15, 1984, P 55-66 25. W.J. BOETTINGER AND S.R. CORIELL, MATER. SCI. ENG., VOL 65, 1984, P 27-36 26. M. CARRARD, M. GREMAUD, M. ZIMMERMANN, AND W. KURZ, ACTA METALL., VOL 40, 1992, P 983-996 27. G.J. DAVIES AND J.G. GARLAND, INT. MET. REV., VOL 20, 1975, P 83-106 28. J. CAMPBELL, INT. MET. REV., VOL 26 (NO. 2), 1981, P 71-108 29. S. KOU, RECENT TRENDS IN WELDING SCIENCE AND TECHNOLOGY, S.A. DAVID AND J.M. VITEK, ED., ASM INTERNATIONAL, 1989, P 137-147 30. S. KOU, WELDING METALLURGY, JOHN WILEY & SONS, 1987, P 161-162 Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

References

1. J.D VERHOEVEN, FUNDAMENTALS OF PHYSICAL METALLURGY, JOHN WILEY & SONS, 1975 2. H.D. SOLOMON, IN TREATISE ON MATERIALS SCIENCE AND TECHNOLOGY, VOL 25, EMBRITTLEMENT OF ENGINEERING ALLOYS, C.L. BRIANT AND S.K. BANERJI, ED., ACADEMIC PRESS INC., 1983, P 525-599 3. J.A. BURTON, R.C. PRIM, AND J. SLICHTER, J. CHEM. PHYS., VOL 21, 1953, P 1987-1991 4. W.A. TILLER, K.A. JACKSON, J.W. RUTTER, AND B. CHALMERS, ACTA METALL., VOL 1, 1953, P 428-437 5. W.A. TILLER AND J.W. RUTTER, CAN. J. PHYS., VOL 34, 1956, P 96-121 6. W.F. SAVAGE, E.F. NIPPES, AND T.W. MILLER, WELD. J., VOL 55, 1976, P 165S-173S 7. F. MATSUDA, T. HASHIMOTO, AND T. SENDA, TRANS. NATL. RES. INST. MET. (JPN), VOL 11, 1969, P 43 8. W.F. SAVAGE, C.D. LUNDIN, AND A.H. ARONSON, WELD. J., VOL 44, 1965, P 420S-425S 9. W.F. SAVAGE, WELD. WORLD, VOL 18, 1980, P 89-113 10. W.F. SAVAGE, C.D. LUNDIN, AND R.J. HRUBEC, WELD. J., VOL 47, 1968, P 420S-425S 11. W.F. SAVAGE, C.D. LUNDIN, AND T.F. CHASE, WELD. J., VOL 47, 1968, P 522S-526S 12. W.F. SAVAGE, E.F. NIPPES, AND J.S. ERICKSON, WELD. J., VOL 55, 1976, P 213S-221S 13. S. KOU AND L. LE, METALL. TRANS. A, VOL 19, 1988, P 1075-1082 14. T. GANAHA, B.P. PEARCE, AND H.W. KERR, METALL. TRANS. A, VOL 11, 1980, P 1351-1359 15. H. NAKAGAWA, M. KATOH, F. MATSUDA, AND T. SENDA, TRANS. JPN. WELD. SOC., VOL 4, 1973, P 11 16. J.A. BROOKS AND M.I. BASKES, WELD MICROSTRUCTURE CHARACTERIZATION AND MODELING, RECENT TRENDS IN WELDING SCIENCE AND TECHNOLOGY, S.A. DAVID AND J.M. VITEK, ED., ASM INTERNATIONAL, 1990, P 93-99 17. H.D. BRODY AND M.C. FLEMINGS, TRANS. AIME, VOL 236, 1966, P 615-624 18. W.W. MULLINS AND R.F. SEKERKA, J. APPL. PHYS., VOL 35, 1964, P 444-451 19. J.S. LANGER AND H. MULLER-KRUMBHAAR, ACTA METALL., VOL 26, 1978, P 1681-1687 20. W. KURZ AND D.J. FISCHER, ACTA METALL., VOL 29, 1981, P 11-20 21. W. KURZ, B. GIOVANOLA, AND R. TRIVEDI, ACTA METALL., VOL 34, 1986, P 823-830 22. R. TRIVEDI AND W. KURZ, ACTA METALL., VOL 34, 1986, P 1663-1670 23. M.J. AZIZ, J. APPL. PHYS., VOL 53, 1982, P 1158-1168 24. W.J. BOETTINGER, D. SHECHTMAN, R.J. SCHAEFFER, AND BIANCANIELLO, METALL. TRANS. A, VOL 15, 1984, P 55-66 25. W.J. BOETTINGER AND S.R. CORIELL, MATER. SCI. ENG., VOL 65, 1984, P 27-36 26. M. CARRARD, M. GREMAUD, M. ZIMMERMANN, AND W. KURZ, ACTA METALL., VOL 40, 1992, P 983-996 27. G.J. DAVIES AND J.G. GARLAND, INT. MET. REV., VOL 20, 1975, P 83-106 28. J. CAMPBELL, INT. MET. REV., VOL 26 (NO. 2), 1981, P 71-108 29. S. KOU, RECENT TRENDS IN WELDING SCIENCE AND TECHNOLOGY, S.A. DAVID AND J.M. VITEK, ED., ASM INTERNATIONAL, 1989, P 137-147 30. S. KOU, WELDING METALLURGY, JOHN WILEY & SONS, 1987, P 161-162 Fundamentals of Weld Solidification Harvey D. Solomon, General Electric Company

Selected References

• CASTING, VOL 15, ASM HANDBOOK, ASM INTERNATIONAL, 1988, P 100-181

• S.A. DAVID AND J.M. VITEK, INT. MATER. REV., VOL 34, 1989, P 213-245 • J.H. DEVLETION AND W.E. WOOD, VOL 6, METALS HANDBOOK, 9TH ED., AMERICAN SOCIETY FOR METALS, 1983, P 21-49 • K. EASTERLING, INTRODUCTION TO THE PHYSICAL METALLURGY OF WELDING, BUTTERWORTHS, 1985 • M.C. FLEMINGS, SOLIDIFICATION PROCESSING, MCGRAW-HILL, 1974 • W. KURZ AND D.J. FISHER, FUNDAMENTALS OF SOLIDIFICATION, 3RD ED., TRANS TECH, 1989 • J.F. LANCASTER, METALLURGY OF WELDING, 3RD ED., GEORGE ALLEN & UNWIN, 1980 • M. RAPPAZ, INT. MATER. REV., VOL 34, 1989, P 93-123 Nature and Behavior of Fluxes Used for Welding D.L. Olson, S. Liu, R.H. Frost, G.R. Edwards, and D.A. Fleming, Colorado School of Mines

Introduction FLUXES are added to the welding environment to improve arc stability, to provide a slag, to add alloying elements, and to refine the weld pool (Ref 1, 2). Different ingredients in the flux system will provide the process with different pyrometallurgical characteristics and thus different weld metal properties (Ref 3, 4). The slag that forms during welding covers the hot weld metal and protects it from the atmosphere. Welding slag consists of the glass-forming components of the flux, as well as inclusions that form in the weld pool, coalesce, rise, and become incorporated into the slag. The need to improve flux formulation to achieve optimal weld metal composition, and ultimately improve the properties of weldments, has led to fundamental studies of weld pyrochemistry. Understanding the thermodynamic and kinetic factors that are prevalent at the electrode, in the arc column, and in the weld pool, has led to more precise prediction of the final weld metal composition (Ref 5, 6, 7, 8, 9, 10, 11, 12, 13, 14, 15).

References

1. G.E. LINNERT, CHAPTER 8, WELDING METALLURGY, VOL 1, AWS, 1965, P 367-396 2. C.E. JACKSON, FLUXES AND SLAGS IN WELDING, WELD. RES. BULL., NO. 190, 1973 3. T. LAU, G.C. WEATHERLY, AND A. MCLEAN, THE SOURCES OF OXYGEN AND NITROGEN CONTAMINATION IN SUBMERGED ARC WELDING USING CAO-AL2O3 BASED FLUXES, WELD. J., VOL 64 (NO. 12), 1985, P 343S-347S 4. T.H. NORTH, H.B. BELL, A. NOWICKI, AND I. CRAIG, SLAG/METAL INTERACTION, OXYGEN, AND TOUGHNESS IN SUBMERGED ARC WELDING, WELD. J., VOL 57 (NO. 3), 1978, P 63S-75S 5. N. CHRISTENSEN AND J. CHIPMAN, SLAG-METAL INTERACTION IN ARC WELDING, WELD. RES. BULL., NO. 15, JAN 1953, P 1-14 6. R.H. FROST, D.L. OLSON, AND S. LIU, PYROCHEMICAL EVALUATION OF WELD METAL INCLUSION EVOLUTION, PROC. 3RD INT. CONF. TRENDS IN WELDING, ASM INTERNATIONAL, JUNE 1992 7. C.A. NATALIE, D.L. OLSON, AND M. BLANDER, PHYSICAL AND CHEMICAL BEHAVIOR OF WELDING FLUXES, ANN. REV. MATER. SCI., VOL 16, 1986, P 389-413 8. J.E. INDACOCHEA, M. BLANDER, N. CHRISTENSEN, AND D.L. OLSON, CHEMICAL REACTIONS WITH FEO-MNO-SIO2 FLUXES, METALL. TRANS. B, VOL 16, 1985, P 237-245 9. U. MITRA AND T.W. EAGAR, SLAG-METAL REACTIONS DURING WELDING, METALL. TRANS. B, VOL 22, 1991, P 65-100

10. C.S. CHAI AND T.W. EAGAR, SLAG-METAL EQUILIBRIUM DURING SUBMERGED ARC WELDING,METALL. TRANS. B, VOL 12, 1981, P 539-547 11. N. CHRISTENSEN, METALLURGICAL ASPECTS OF ARC WELDING, WELD. J., VOL 27, 1949, P 373S-380S 12. O. GRONG, D.L. OLSON, AND N. CHRISTENSEN, CARBON OXIDATION IN HYPERBARIC MMA WELDING, MET. CONSTRUCT., VOL 17. DEC 1985, P 810R-814R 13. U. MITRA AND T.W. EAGAR, SLAG-METAL REACTIONS DURING SUBMERGED ARC WELDING OF ALLOY STEELS, METALL. TRANS. A, VOL 15, 1984, P 217-227 14. T.W. EAGAR, SOURCES OF WELD METAL OXYGEN CONTAMINATION DURING SUBMERGED ARC WELDING, WELD. J.,VOL 57, 1978, P 76S-80S 15. N. CHRISTENSEN, WELDING METALLURGY, LECTURE NOTES, NTH, 1979 Nature and Behavior of Fluxes Used for Welding D.L. Olson, S. Liu, R.H. Frost, G.R. Edwards, and D.A. Fleming, Colorado School of Mines

Equilibrium Parameters Equilibrium is not achieved during welding, because of the very large temperature and density gradients, the short reaction times, and the large electric currents. Despite these expected departures from equilibrium, thermodynamic considerations can be used as a guide for constraining chemical reactions and mechanisms involved in welding (Ref 5, 7, 9). A common approach is to assume that extremely high temperatures and high surface-to-volume ratios allow thermodynamic equilibrium to be locally attained, in spite of the short reaction times available. Further complicating the issue is the dependence on welding parameters of the chemical partitioning between the slag and the weld metal (Ref 3, 7). This dependence suggests that the pyrometallurgical reactions involved are influenced by the processes that occur at the electrode tip. As a specific example, the welding parameters affect metal-droplet size, which in turn alters the chemical kinetics. Nonetheless, thermodynamics will reveal the direction taken by the chemical reactions (but will not accurately predict weld metal composition). The effective slag-metal reaction temperature has been estimated (by pyrochemical analysis of slag-metal compositions) to be approximately 1900 °C (3450 °F), a temperature intermediate between that of the hot spot at the arc root (2300 °C or 4170 °F) and the melting point of iron (1500 °C or 2730 °F). Thus, during the droplet lifetime, the average temperature is effectively 1900 °C (2730 °F). During this period, oxygen from the hot spot reactions is distributed throughout the droplet, and that oxygen reacts with the metallic elements to form oxides; these products of oxidation pass into the slag, and the slag-metal reactions tend toward equilibrium. Estimates of the time for which the molten slag and molten metal are in contact range from 3 to 8 s. During the process, the gaseous phase in the arc cavity contacts the metal for an estimated 0.5 to 1.0 s. Effect of Oxygen The most important chemical reagent in controlling weld metal composition, and thus microstructure and properties, is oxygen (Ref 3, 4, 14, 16, 17, 18, 19). Oxygen directly reacts with alloying elements to alter their effective role by: • • •

REDUCING HARDENABILITY PROMOTING POROSITY PRODUCING INCLUSIONS.

All three effects are significant to weld quality. Oxygen is introduced into the weld pool at high temperatures by:

• • • •

THE PRESENCE OF OXIDE FLUXES THAT DISSOCIATE IN THE ARC THE SLAG-METAL REACTIONS IN THE WELD POOL THE OXIDES ON THE SURFACE OF BAKED METALLIC POWDERS MIXED WITH FLUX OR ON ELECTRODE THE ASPIRATION OF ATMOSPHERE (AIR) INTO THE ARC.

Shielding gas may contain oxidizing reagents; a common gas is 75Ar-25CO2 or 100% CO2. As a 100 to 1000 ppm, depending on the type of welding consumable used. The weld metal oxygen measured directly at the molten electrode tip has been reported to be as high as 1400 ppm. Individual droplets have been found to contain as much as 2000 ppm O2. There are two views concerning the genesis of the high oxygen content. One suggests that pyrochemical or electrochemical reactions (or both) provide oxygen to the electrode tip, and then further oxidation occurs within the droplet as it passes through the arc. The other view is that the high oxygen levels of the droplet represent the maximum buildup of oxygen in the electrode prior to detachment, with limited reaction during flight across the arc to the weld pool. In either case, high oxygen concentrations are introduced into the weld pool by the welding process. At the melting point, the solubility of oxygen in pure liquid iron is approximately 1600 ppm at 100 kPa (1 atm) pressure. During solidification, solubility decreases to about 860 ppm at 1500 °C (2730 °F) in δ-Fe. Most of the alloying elements present in liquid steel reduce oxygen solubility through deoxidation equilibria. Steelmaking processes typically yield analytical oxygen levels ranging from 70 to 100 ppm. Welds typically pick up oxygen to levels of several hundred ppm, then deoxidize to oxygen levels of around 200 to 300 ppm. Deoxidation of the weld metal occurs in two separate steps, the first being the primary deoxidation of the weld pool (Ref 6). Secondary deoxidation occurs during solidification as solute concentrations increase within the intercellular or interdendritic regions. The secondary deoxidation will either form very small inclusions or will coat the interdendritically trapped primary inclusions. The high oxygen concentrations added to the weld pool by the metal droplets significantly affect deoxidation. Figure 1 shows experimentally measured soluble oxygen concentrations for various deoxidants (solid lines), along with deoxidation curves predicted by the solubility products (broken lines). The experimental deviation is caused by interactions of the deoxidant with other alloy elements. If the oxygen and metallic element concentrations resulting from the welding process exceed the equilibrium concentration for a specific reaction, inclusions will result. The ability to form a specific inclusion will correspond directly to the position of the weld pool composition relative to the activity plot for this inclusion. Thus, the thermodynamic order for the formation of primary oxides would be: Al2O3 > Ti2O3 > SiO2 > Cr2O4 > MnO.

FIG. 1 DEOXIDATION EQUILIBRIA IN LIQUID IRON ALLOYS AT 1600 °C (2910 °F). THE BROKEN LINES SHOW DEOXIDATION EQUILIBRIA PREDICTED BY SOLUBILITY PRODUCT CALCULATIONS. THE SOLID LINES SHOW EXPERIMENTALLY DETERMINED SOLUBLE OXYGEN CONCENTRATIONS FOR VARIOUS DEOXIDANTS. THE EXPERIMENTAL DEVIATION IS CAUSED BY VARIATIONS IN THE ACTIVITY COEFFICIENTS WITH INCREASING DEOXIDANT CONCENTRATION. SOURCE: REF 20

Inclusion Formation. Inclusions form as a result of reactions between metallic alloy elements and nonmetallic tramp elements, or by mechanical entrapment of nonmetallic slag or refractory particles. Inclusions may include: • • • • • •

OXIDES SULFIDES NITRIDES CARBIDES OTHER COMPOUNDS MULTIPLE PHASES

Among these, oxides and complex oxides occur most frequently in the size range known to influence weld metal microstructure.

Using only thermodynamic considerations in the analysis of slag-metal reactions, the following reactions may be considered as the ones that describe inclusion formation:

X M + Y O = (MXOY)

(EQ 1)

where the underlining of a component M ( M ) and the component O ( O ) indicate that M and O, respectively, are dissolved in the metal, and:

X M + Y(FEO) = (MXOY) + YFE

(EQ 2)

The equilibrium constants (K1 and K2) for the above two reactions are: K1 =

( AM x Oy ) [aM ]x [ao ] y

(EQ 3)

where ai is the activity, a function of the concentration, for component i, and: K2 =

( AM x Oy ) [aM ]x [aFeO ] y

(EQ 4)

Equilibrium compositions for the weld deposit can be estimated with Eq 1, 2, 3, and 4 and can be used to predict trends for the weld pool pyrochemical reactions. The actual compositions may, however, differ from the calculated values due to alloying-element partitioning during cellular or dendritic solidification, commonly observed in steel weldments. During solidification of a weld metal, solute elements segregate to the liquid at the solid/liquid interface, and the solute concentrations can reach high levels in the interdendritic regions, as suggested in Fig. 2. Neglecting solid diffusion, the solute composition in the liquid at the solid/liquid interface can be modeled (Ref 6) by the nonequilibrium lever rule, or Scheil Equation:

CL = C0FLK-1

(EQ 5)

where C0 is the bulk concentration in the weld pool, CL is the solute concentration in the liquid at the interface, and k is the equilibrium partition ratio.

FIG. 2 SCHEMATIC SHOWING THE SOLID AND LIQUID COMPOSITION PROFILES MODELED BY EQ 5 BASED ON THE ASSUMPTION THAT THERE IS COMPLETE LIQUID DIFFUSION AND NO SOLID DIFFUSION. C0, THE INITIAL ALLOY COMPOSITION; K, THE PARTITION RATIO OF THE SOLID TO LIQUID COMPOSITIONS ON THE EQUILIBRIUM PHASE DIAGRAM; KC0, THE INITIAL COMPOSITION OF THE SOLID; AND CS AND CL, THE SOLID AND LIQUID COMPOSITIONS, RESPECTIVELY, AT THE SOLIDIFICATION INTERFACE. SOURCE: REF 4

The equilibrium partition ratio, k, controls the direction and the extent of segregation. For most alloy elements in steel, the partition ratio is less than one, and the element segregates to the interdendritic liquid. Consider the deoxidation equilibrium as represented by the dissolution reaction for a complex oxide:

MXNYOZ = X M + Y N + Z O

(EQ 6)

The free energy change associated with the dissolution of the MxNyOz can be written as: ∆G = ∆G° + RT ln

[ M ]x [ N ] y [O]Z [ M x N y Oz ]

(EQ 7)

where [M], [N], and [O] are the solute activities in the liquid; and x, y, and z are the stoichiometric constants from Eq 6. The value [MxNyOz] is the activity for the specific inclusion and can be assumed to have the value of one. The ratio of the activities of the reactants to that of the oxide can be termed the activity quotient (Q): Q=

[ M ]x [ N ] y [O]Z [ M x N y Oz ]

(EQ 8)

With Eq 5 representing the extent of segregation, the solute activities in the interdendritic liquid can be written as:

[M]=[

M][MO]

[N]=[

N][N0]

[O] = [

O][O0]

(EQ 9A) (EQ 9B) (EQ 9C)

In Eq 9a, 9b, and 9c, [MO], [NO], and [OO] are the bulk concentrations of M, N, and O in the melt; and [γM], [γN], and [γO] are the activity coefficients for the solutes. Substituting Eq 9a, 9b, and 9c into Eq 8 gives the free energy as a function of the remaining liquid fraction:

(EQ 10)

Equation 10 expresses the free energy driving force for oxide dissolution. The first term on the right side, ∆G, represents equilibrium conditions. The second term represents the departure from equilibrium caused by changes in reactant concentrations or by segregation during solidification. At equilibrium, the free energy driving force, ∆G, is zero, and the equilibrium concentrations can be found by equating the two terms on the right-hand side of Eq 10. At equilibrium, the activity quotient becomes the equilibrium constant, Keq:

(EQ 11) Inclusion precipitation is possible when the concentrations of oxygen and deoxidants exceed the equilibrium values for a particular oxide. This condition can be expressed with the ratio of the activity quotient to the equilibrium coefficient:

(EQ 12) A precipitation index less than unity indicates that the concentrations of oxygen and deoxidants are below the equilibrium value, and the precipitation of the inclusions will not occur. A value greater than 1.0 indicates that concentrations are sufficiently high for precipitation according to the methodology described above. The compounds Al2O3, Ti2O3, and SiO2 are some of the oxides that will form in a low-carbon low-alloy steel weld. Multiple reactions can occur, and different oxides may appear in the same weld when more than one of these deoxidizers is present. Metal Transferability During Pyrochemical Reactions. The final weld metal concentration for a particular element is

made up of contributions from the filler wire, flux, and base metal; however, losses caused by the welding process vary for each element. Delta Quantity. The nominal composition of each weld can be calculated considering just the dilution effect of the filler

wire and base metal. The extent of loss or transfer of a specific element can be evaluated by a quantity, which expresses the difference between the analytical and the nominal composition. These quantities, designated delta quantity in this article, indicate the effect of the flux on element transfer during welding. A positive delta quantity indicates an elemental transfer from the flux to the weld metal. A negative delta quantity suggests an elemental loss from the weld pool to the slag. A null delta quantity for a specific element suggests an apparent equilibrium condition, in which the flux and slag content for that element are the same. Flux compositions with null delta-quantity behavior have been used to make equilibrium calculations and thus achieve a better understanding of the chemical reactions involved in welding (Ref 8). Investigators have also quantified elemental transfer by measuring similar neutral points (null delta quantity) for various flux systems and have developed a thermodynamic model capable of predicting neutral points for some slag systems (Ref 10). Arc Stabilizers. Arc welding fluxes are compositionally more complex than fluxes used in other metallurgical processes,

such as steelmaking. Many of the additions to the flux are not designed to assist weld metal refinement. Some additions are present to promote arc stability, generate plasma and protective gases, control viscosity, support out-of-position welding, and promote slag detachability. The welding arc requires an inert or chemically reducing plasma and shielding gas that can be easily ionized. Additions must be made to the flux to achieve the necessary current-carrying capacity and to maintain arc stability. Specific additions will be necessary for the various modes of current (direct current, dc, or alternating current, ac) and polarity. Alkali metal, zirconium, and titanium additions to the arc affect the ionization process and the ease with which the welding is reinitiated (reinitiation is required 50 to 60 times per second with ac welding). These additions come to the welding flux as feldspar (alkali aluminum silicates), rutile, lithium carbonate, titanium aluminate, and potassium oxalate, and they play an especially important role when welding either in the dc electrode negative (DCEN) mode or in the alternate current (ac) mode. Arc stabilizers are also important in high-speed welding when the cathode and anode spots are less stable. These additions are part of the AWS classification of electrodes, as seen by the fourth digit (E XXX X ) in the classification standard for steel electrodes for shielded metal arc welding (SMAW) given in Table 1. The electrodes designed to perform with ac mode or in the DCEN mode current have been specified as containing high titania or alkali metal (potassium or sodium) additions. Special additions such as Li2O have been used to achieve multipurpose results-for example, reducing the viscosity while increasing the arc stability.

TABLE 1 AWS CLASSIFICATION OF SELECTED SMAW ELECTRODES FOR WELDING MILD AND LOWALLOY STEELS

ELECTRODE DESIGNATION EXX10 EXXX1 EXXX2 EXXX3 EXXX4

CURRENT AND POLARITY(A) DCEP AC, DCEP AC, DCEN AC, DCEP, DCEN AC, DCEP, DCEN

PENETRATION ARC STABILIZERS DEEP DEEP MEDIUM LIGHT LIGHT

HIGH CELLULOSE-NA HIGH CELLULOSE-K HIGH RUTILE-NA HIGH RUTILE-K RUTILE-IRON POWDER

EXX24 EXXX5 EXXX6 EXX27

AC, DCEP, DCEN DCEP AC, DCEP AC, DCEN, DCEP

LIGHT MEDIUM MEDIUM MEDIUM

EXXX8

AC, DCEP

MEDIUM

EXX28

AC, DCEP

MEDIUM

RUTILE-IRON POWDER LOW H-NA LOW H-K IRON OXIDE-IRON POWDER LOW HYDROGEN-IRON POWDER LOW HYDROGEN-IRON POWDER

% FE 0-10 0 0-10 0-10 2540 50 0 0 50 2540 50

(A) DCEP, DC ELECTRODE POSITIVE, DCEN, DC ELECTRODE NEGATIVE Changes in Flux Composition With Delta Quantity. Figure 3 shows a transfer of manganese for welds made with

SiO2-TiO2-CaO-1Na2O flux at constant SiO2 content in the flux. These values vary widely, depending on other alloy concentrations. The data imply that manganese is almost always lost to the slag, and that the activity of manganese varies considerably with changes in the amount of titania in the flux. Manganese is very important to weld metal hardenability and must be closely controlled to obtain the optimum weld microstructure. Controlling the weld metal manganese concentration in the titania-containing flux systems would require strict compositional control of the welding flux to ensure the correct manganese concentration in the weld metal.

FIG. 3 PLOT OF CHANGES IN MANGANESE CONTENT IN THE WELD VERSUS THE VARIATION IN FLUX COMPOSITION AS A FUNCTION OF TWO-HEAT INPUTS. THE SIO2 CONTENT WAS MAINTAINED AT 40% THROUGHOUT THE FLUX-CORED ARC WELDING PROCESS. SOURCE: REF 21

In the same system, the delta titanium showed a constant increase with increasing titania content of the flux (Fig. 4). Because of the large amounts of titania in the flux, it is not surprising that the welds show positive compositional deviations for titanium. These results also suggest that control of the weld metal titanium content will require tight control of the limits of the TiO2 content of the flux.

FIG. 4 PLOT OF CHANGES IN TITANIUM CONTENT IN THE WELD VERSUS THE VARIATION IN FLUX COMPOSITION AS A FUNCTION OF TWO HEAT INPUTS. THE SIO2 CONTENT WAS MAINTAINED AT 40% THROUGHOUT THE FLUX-CORED ARC WELDING PROCESS. SOURCE: REF 21

If the delta quantity changes rapidly with flux composition, it may be difficult to maintain specific weld metal composition, microstructure, and properties with variation in flux composition. The magnitude of the delta quantity for specific elements is most often not as serious a concern as a rapid change in the delta quantity with variations in flux chemistry and welding parameters. The magnitude of the delta quantity can be adjusted by altering the alloy content of the welding wire or the amount of ferro-additions to the flux. Thus, a suitable combination of wire, flux, and welding parameters should achieve negligibly small delta quantities of major alloying elements. Shielding Gas. When the shielding gas that protects the weld pool comes from the flux, it is necessary to understand the

decomposition of specific flux components. Two common shielding gas atmospheres from flux dissociation are hydrogen and CO/CO2. The hydrogen gas can be produced by the decomposition of cellulose (wood flour or similar hydrocarbons). A CO/CO2 atmosphere results from the decomposition of carbonates such as limestone (CaCO3) or dolomite [CaMg(CO3)2]. The CO/CO2 atmosphere can be balanced to provide a reducing (and thus protective) atmosphere. At high temperatures, CO2 or CO will react with carbon to achieve equilibrium, which requires the presence of CO and CO2. It is the relative amounts of CO and CO2 that determine the reducing/oxidizing nature of the arc environment. The CO/CO2 ratio also determines the recovery of specific alloying elements. Typical plasma atmospheres for both hydrogentype and low-hydrogen-type electrodes are given in Table 2.

TABLE 2 GAS COMPOSITION OF WELDING ARC OBTAINED FROM SPECIFIC TYPES OF WELDING ELECTRODES

AWS TYPE COMPOSITION, WT% DESIGNATION H2O H2 CO2 CO 6010 CELLULOSIC 16 41 3 40

6015

BASIC

2

2

19

77

Source: Ref 15

In submerged arc welding, the covering flux also produces the protective shielding gas and plasma. Other gaseous phases, including flourine-bearing components, are also part of the plasma. The effect of arc environment (for example, amounts of CO and CO2) on weld metal chemistry control for hyperbaric welding with a basic electrode is shown in Fig. 5. The high-pressure welding allows evaluation of the CO reaction. Considering the CO reaction:

C + O = CO

(EQ 13)

the law of mass action gives:

(EQ 14) where [C] and [O] are the weld metal carbon and oxygen contents, respectively. At equilibrium the partial pressure of CO is directly related to the total pressure by Dalton's Law.

FIG. 5 EFFECT OF PRESSURE ON THE PRODUCT M = [%C][%O] FOR HYPERBARIC WELDING WITH A BASIC ELECTRODE. SOURCE: REF 12

In Fig. 5, two lines are indicated. The solid line plots the product of the weld metal oxygen and carbon as a function of total pressure when the analytical weld metal oxygen concentration, [%O]anal, was used. The broken line indicates a similar relationship; however, in this case, the weld metal oxygen concentration has been corrected for the displacement of oxygen because of the formation of manganese silicate inclusions during welding and its transport to the slag. The equilibrium oxygen content, [%O]eq, in the liquid steel at high temperatures is then given by the following expression:

(EQ 15) where ∆[%Si] and ∆[%Mn] represent the difference between the expected compositions of these elements (from knowledge of initial consumables and base plate compositions as well as dilution) and the actual compositions measured. The linearity of these lines in Fig. 5 is evidence of the strong influence of the CO reaction in arc welding. The evidence of water reaction control for some electrodes (for example, cellulosic electrodes) can be seen in Fig. 6. With an increase in the degree of oxidation of the deposited metal, the weld metal hydrogen content decreases (Ref 22). The relationship between weld metal oxygen and hydrogen contents as seen in Fig. 6 is consistent with the functional form expected from the law of mass action for a H2O reaction. It is apparent that with increasing weld metal oxygen content there is a significant reduction in weld metal hydrogen. The hydrogen content of the deposited metal can be reduced by increasing the CaCO3 content of the coating. With the ever-increasing requirements of weld-deposit properties and the increasing need for higher productivity, pyrochemistry will play an important role for flux formulation and weld-property prediction.

FIG. 6 PLOT OF WELD METAL OXYGEN CONTENT VERSUS WELD METAL HYDROGEN CONTENT WHEN WELDING WITH ELECTRODES THAT CONTAIN CHROMIUM AND NIOBIUM IN THEIR COATINGS. SOURCE: REF 22

Basicity Index As indicated in the section "Metal Transferability During Pyrochemical Reactions" in this article, the transfer of alloying elements during welding depends strongly on the physical and chemical properties of the flux. The ability to correlate flux properties with weld metal chemical compositions and properties is essential to understand the interactions between weld metal and flux. Because of the incomplete understanding of the thermodynamic properties of slags, the empirical concept of basicity has been applied to predict flux and weld properties. A basicity index, BI, for welding has been proposed:

(EQ 16)

where the chemical components are given in wt%. When the basicity index for a given flux is less than one, the flux is considered acidic. A flux with a BI between 1.0 and 1.2 can be classified as neutral. A flux with a BI greater than 1.2 is considered basic. In general, the higher the basicity, the cleaner the weld with respect to nonmetallic inclusions (that is, lower weld metal oxygen content) (Ref 23). Figure 7 illustrates the correlation between weld metal oxygen and the basicity index for some flux systems. The weld metal oxygen content drops significantly as the BI is increased to 1.2 and then remains relatively constant at about 250 ppm O. The correlation between weld metal oxygen, which is an indication of weld metal toughness, and basicity is acceptable for some welding flux systems, especially those that are primarily based on CaO, MgO, and SiO2. However, Eq 16 cannot be used to correlate the strength and toughness of welds made with high flux concentrations of Al2O3, TiO2, ZrO2, MnO, FeO, and CaF2. Although numerous basicity formulas have been considered, none has been flexible enough to deal with high amphoteric oxide contents. There is still some concern over utilizing this index (which was primarily conceived by steelmakers for evaluating sulfur refinement) for predicting weld metal oxygen, or for use as a general criterion for weld metal quality (Ref 7). Additionally, the BI does not consider the physical properties of the fluxes, nor does it explain the kinetics of the flux-metal reactions.

FIG. 7 EFFECT OF WELD METAL OXYGEN CONTENT ON FLUX BASICITY INDEX WHEN USING THE SUBMERGED ARC WELDING PROCESS. SOURCE: REF 23

Basicity index has been used by the welding industry as a measure of expected weld metal cleanliness and mechanical properties. Consequently, manufacturers of welding consumables have classified and advertised their fluxes with this index. It is believed that high basicity means high toughness, a quality of great interest to the engineer, while an acidic flux means excellent slag behavior, a characteristic of interest to the welder attempting to improve weld bead morphology and deposition rate. Table 3 classifies the various coating formulations for SMAW electrodes using the descriptors of cellulosic, basic, acid, and rutile (alternating current stabilizer and slag former).

TABLE 3 ELECTRODE COATING FORMULATIONS OF SELECTED SMAW ELECTRODES

Electrode Type

Coating formulation(a) AWS designation

Comments

Rutile

Cellulose

Quartz

Carbonates

Ferromanganese

Organics

Iron oremanganese ore

Calcium carbonate

Complex silicates

Fluorspar

Ferroalloys

Cellulosic 6010

2060

10-50

1530

0-15

5-10

...

...

...

...

...

...

Rutile

6012 6013

4060 2040

... ...

1525 1525

0-15 5-25

10-12 1.4-14

2-6 0-5

... ...

... ...

... ...

... ...

... ...

Acidore(a)

6020

...

...

X

X

X

...

X

...

X

...

...

Basic

7015

010

...

0-5

...

...

...

...

20-50

...

20-40

5-10

Source: Ref 24

(A)

X, DATA UNAVAILABLE.

Cellulose promotes gas shielding in the arc region. Hydrogen increases heat at weld. High hydrogen content (30-200 ppm). Deep penetration, fast cooling weld Slags mainly for slag shielding relatively high hydrogen content (15-30 ppm). High inclusion content in weld deposit Relatively high hydrogen content. High slag content in weld metal Relatively low hydrogen levels ( 10 ppm), hence commonly used in welding low-alloy construction steels. Electrodes should be kept dry. Low inclusion content in weld deposit

Pyrochemical Kinetics During Welding The ability of a flux to refine as well as protect the weld pool is related to the mass transport processes in the flux. The flux should melt approximately 200 °C (360 °F) below that of the alloy for proper flux coverage and for protection of the weld deposit. One of the most important physical properties of a flux is its slag viscosity, which not only governs the way the slag flows and covers the molten weld pool but also strongly affects the transport processes involved in pore removal, deoxidation, and retention of alloying additions. The chemical processing and refining by the flux to achieve a weld deposit with low concentrations of oxygen and sulfur and optimal concentration of hardenability agents (carbon, manganese, chromium, molybdenum, nickel, and so on) may not be achieved unless slag viscosity is also adequate. The viscosity is strongly temperature-dependent, so the use of various heat inputs during welding may require different flux compositions to produce the matching slag viscosity. For simple ionic melts, the viscosity has been shown to follow an Arrhenius-type temperature dependence:

(EQ 17) where ηis the viscosity, η0 is a system constant, Eη is the viscous activation energy, RT is the gas constant, and T is the temperature. For polymeric melts, Eq 17 often does not hold true. It has been shown that in certain cases a modified equation can be used:

(EQ 18)

where T0 is a constant for a given flux composition. The presence of entrapped particulates tends to increase the slag viscosity. Slag viscosity is also affected by composition. The compositional dependence is commonly reported by considering Eη to be a function of composition at constant temperature. The slag must be fluid enough so that it flows and covers the molten weld pool but must be viscous enough so that it does not run away from the molten metal and flow in front of the arc, leading to possible overlapping by the weld metal. (For overhead welding, surface tension becomes a primary factor because fluidity reduces coverage [opposite gravitational vector].) It has been reported that if the manganese silicate flux viscosity at 1450 °C (2640 °F) is above 0.7 Pa · s (7 P), a definite increase in weld surface pocking will occur. Pock marks have been associated with easily reducible oxides in the flux, which contribute oxygen to the weld pool. The weld pool reacts with carbon to form carbon monoxide, which cannot be transported through a high-viscosity flux and is trapped at the liquid-metal/flux interface. The result is a weld metal surface blemished by surface defects or pocks. Because viscosity is sensitive to temperature and thus heat input, pocking can be the evidence that a flux formulated for high-current welding is being used at too low a current or too great a travel speed. The viscosity of most welding fluxes at 1400 °C (2550 °F) is in the range of 0.2 to 0.7 Pa · s (2 to 7 P). Slag viscosity also affects the shape of the weld deposit and must be carefully controlled when covered electrodes are used out of position. The higher the slag viscosity, the greater the weld penetration in submerged arc welding. However, this benefit must be balanced, because if the viscosity is too high, the gaseous products cannot escape the weld pool, resulting in unacceptable porosity. This condition can be monitored by observing the density of pores trapped in the underside of the detached slag. Detached slags manifesting a honeycomb structure suggest a severe weld metal porosity problem. This condition usually means that a given flux has experienced an insufficient heat input for the effective transport of gas through the slag.

References cited in this section

3. T. LAU, G.C. WEATHERLY, AND A. MCLEAN, THE SOURCES OF OXYGEN AND NITROGEN CONTAMINATION IN SUBMERGED ARC WELDING USING CAO-AL2O3 BASED FLUXES, WELD. J., VOL 64 (NO. 12), 1985, P 343S-347S

4. T.H. NORTH, H.B. BELL, A. NOWICKI, AND I. CRAIG, SLAG/METAL INTERACTION, OXYGEN, AND TOUGHNESS IN SUBMERGED ARC WELDING, WELD. J., VOL 57 (NO. 3), 1978, P 63S-75S 5. N. CHRISTENSEN AND J. CHIPMAN, SLAG-METAL INTERACTION IN ARC WELDING, WELD. RES. BULL., NO. 15, JAN 1953, P 1-14 6. R.H. FROST, D.L. OLSON, AND S. LIU, PYROCHEMICAL EVALUATION OF WELD METAL INCLUSION EVOLUTION, PROC. 3RD INT. CONF. TRENDS IN WELDING, ASM INTERNATIONAL, JUNE 1992 7. C.A. NATALIE, D.L. OLSON, AND M. BLANDER, PHYSICAL AND CHEMICAL BEHAVIOR OF WELDING FLUXES, ANN. REV. MATER. SCI., VOL 16, 1986, P 389-413 8. J.E. INDACOCHEA, M. BLANDER, N. CHRISTENSEN, AND D.L. OLSON, CHEMICAL REACTIONS WITH FEO-MNO-SIO2 FLUXES, METALL. TRANS. B, VOL 16, 1985, P 237-245 9. U. MITRA AND T.W. EAGAR, SLAG-METAL REACTIONS DURING WELDING, METALL. TRANS. B, VOL 22, 1991, P 65-100 10. C.S. CHAI AND T.W. EAGAR, SLAG-METAL EQUILIBRIUM DURING SUBMERGED ARC WELDING,METALL. TRANS. B, VOL 12, 1981, P 539-547 12. O. GRONG, D.L. OLSON, AND N. CHRISTENSEN, CARBON OXIDATION IN HYPERBARIC MMA WELDING, MET. CONSTRUCT., VOL 17. DEC 1985, P 810R-814R 14. T.W. EAGAR, SOURCES OF WELD METAL OXYGEN CONTAMINATION DURING SUBMERGED ARC WELDING, WELD. J.,VOL 57, 1978, P 76S-80S 15. N. CHRISTENSEN, WELDING METALLURGY, LECTURE NOTES, NTH, 1979 16. O. GRONG, T.A. SIEWERT, T.A. MARTINS, AND D.L. OLSON, A MODEL FOR THE SILICONMANGANESE DEOXIDATION OF STEEL WELD METAL, METALL. TRANS. A, VOL 17 (NO. 10), 1985, P 1797-1807 17. T.H. NORTH, THE DISTRIBUTION OF MANGANESE BETWEEN SLAG AND METAL DURING SUBMERGED ARC WELDING, WELD. RES. ABROAD, VOL 23 (NO. 1), 1977, P 2-40 18. T. BONISZEWSKI, BASIC FLUXES AND DEOXIDATION IN SUBMERGED ARC WELDING OF STEEL, METAL. CONSTR. BRIT. WELD. J., VOL 6, 1974, P 128 19. L. DAVIS, AN INTRODUCTION TO WELDING FLUXES FOR MILD AND LOW ALLOY STEELS, WELDING INSTITUTE, 1981 20. MAKING, SHAPING AND TREATING OF STEEL, UNITED STATES STEEL CORPORATION, 1984 21. P.S. DUNN, C.A. NATALIE, AND D.L. OLSON, SOL-GEL FLUXES FOR FLUX CORED WELDING CONSUMABLES, J. MATER. ENERGY SYSTEMS, VOL 8 (NO. 2), 1986, P 176-184 22. L.I. SOROKIN AND Z.A. SIDLIN, THE EFFECT OF ALLOYING ELEMENTS AND OF MARBLE IN AN ELECTRODE COATING ON THE SUSCEPTIBILITY OF A DEPOSITED NICKEL CHROME METAL TO PORE FORMATION, SVAR. PROIZVOD., NO. 11, 1974, P 7-9 23. S.S. TULIANI, T. BONISZEWSKI, AND N.F. EATON, NOTCH TOUGHNESS OF COMMERCIAL SUBMERGED ARC WELD METAL, WELD. MET. FABR., VOL 37 (NO. 8), 1969, P 27 24. T.G.F. GRAY, J. SPENCE, AND T.H. NORTH, RATIONAL WELDING DESIGN, BUTTERWORTHS, 1975 Nature and Behavior of Fluxes Used for Welding D.L. Olson, S. Liu, R.H. Frost, G.R. Edwards, and D.A. Fleming, Colorado School of Mines

Alloy Modification Other additions to the welding consumables are required to make alloy additions to the weld pool, usually in the form of powder metal or ferro-additions (Ref 1). Often the composition of the wire that makes up the rod for the shielded metal arc electrode from a specific manufacturer is the same, regardless of the alloy to be welded. Alloying is achieved by

powder metal additions to the flux coating. Manganese, silicon, chromium, niobium, and other alloying additions are adjusted in the weld pool by ferroalloy powder additions. Specially prepared alloy additions of Fe-50Si, Fe-80Mn, Fe60Mn-30Si, and others are used. One concern in formulating electrodes using ferro-additions is the alloying element recovery (that is, the amount of the element that is transferred across the arc and into the weld deposit). Values for the recovery of typical elements in steel welding are given in Table 4. The metal losses are either to the slag or to the fume.

TABLE 4 RECOVERY OF ELEMENTS FROM SELECTED ELECTRODE COVERINGS

ALLOY ELEMENT FORM OF MATERIAL IN ELECTRODE COVERING ALUMINUM FERROALUMINUM BORON FERROBORON CARBON GRAPHITE CHROMIUM FERROCHROMIUM NIOBIUM FERROCOLUMBIUM COPPER COPPER METAL MANGANESE FERROMANGANESE MOLYBDENUM FERROMOLYBDENUM NICKEL ELECTROLYTIC NICKEL SILICON FERROSILICON TITANIUM FERROTITANIUM VANADIUM FERROVANADIUM

APPROXIMATE RECOVERY OF ELEMENT, WT% 20 2 75 95 70 100 75 97 100 45 5 80

Source: Ref 1 Slipping and Binding Agents. In the case of SMAW electrodes, slipping agents are also added to the green flux formulation to improve the extrudability of the flux onto the rod (Ref 1). Glycerin, china clay, kaolin, talc, bentonite, and mica have all been used as slipping agents. Binding agents can be classified into two types. The first type comprises binders that bond the flux components to the rod without introducing a hydrogen source. These low-hydrogen binders include sodium silicate and potassium silicate. The second type of binding agent does function as a hydrogen source. The binders used for high-hydrogen electrodes can be organic in nature and include gum arabic, sugar, dextrine, and other specialized synthetic organic binders. Slag Formation. Slag, a mixture of glass and crystalline structure, must solidify on the already solidified weld deposit to

protect the surface from oxidation during cooling. Specific physical properties are required of the slag. It must melt below the melting temperature of steel (~ 1450 °C, or 2640 °F), must have a density significantly less than steel to reduce slag entrapment in the weld deposit, must possess the proper viscosity in the temperature range of 1450 to 1550 °C (2640 to 2820 °F), and must easily detach from the weld deposit after welding. Silicates, aluminates, and titanates are all primary slag formers. The high-valence cations of these compounds produce a bonding network that can promote glass formation. Most electrodes produce silicate or titanate slag. Silicates of such elements as manganese produce very workable glass deposits, but also produce weld deposits that are relatively high in weld metal oxygen content. These fluxes are said to range from acidic to neutral. The titanate and aluminate fluxes produce more rigid and stable oxide covers. The result is a lower concentration of weld metal oxygen, but a higherviscosity slag. Aluminates and calcium-bearing compounds are common additions to basic fluxes for the submerged arc welding of high-toughness linepipe steels. Minerals used for slag formation include: • • • • •

RUTILE POTASSIUM TITANATE ILMENITE ALUMINA SILICA FLOUR

• • • •

IRON POWDER FLUORSPAR FELDSPAR MANGANESE DIOXIDE

Asbestos (up to 50%) was used as a slag former, but has been phased out by welding consumable manufacturers. Slag detachability is a serious productivity concern for steel fabricators. The relative ease of slag detachability influences

the economic advantages of flux-related welding processes. Residual slag on the weld deposit promotes slag stringers in multipass weldments, limits the effective use of narrow-gap flux-related welding processes, and may reduce the corrosion resistance of the weldment. Welding flux formulators have modified flux compositions to alleviate or reduce this hindrance (Ref 25). Poor slag removal has been reported to occur when the flux contains fluorite. Slags containing spinels generally have been found to attach tenaciously to the weld deposit. Slags with cordierite and (Cr,Mn,Mg)O · (Cr,Mn,Al)2O3 type spinel phases have been reported to be difficult to remove from stainless steel weldments. It has also been reported to be difficult to remove from stainless steel weldments. It has also been reported that if (CaO)2SiO2, Cr2TiO5, and FeTiO5 are present, the slag readily detaches from the weld deposit (Ref 25). The compositional range for acceptable slag detachability has been reported for the CaO · CaF2 · SiO2 system and the CaO · TiO2 · SiO2 system. Easier slag removal has also been related to deoxidation with aluminum instead of titanium. Increasing the Al2O3 content in the flux has demonstrated improved detachability (Ref 25). Differences in the coefficient of thermal expansion between the slag and the weld are also important.

References cited in this section

1. G.E. LINNERT, CHAPTER 8, WELDING METALLURGY, VOL 1, AWS, 1965, P 367-396 25. D.L. OLSON, G.R. EDWARDS, AND S.K. MARYA, THE PHYSICAL AND CHEMICAL BEHAVIOR ASSOCIATED WITH SLAG DETACHABILITY DURING WELDING, FERROUS ALLOY WELDMENTS, VOL 67-70, TRANS TECH, 1992, P 253-268

Nature and Behavior of Fluxes Used for Welding D.L. Olson, S. Liu, R.H. Frost, G.R. Edwards, and D.A. Fleming, Colorado School of Mines

Types of Fluxes Producing a weld deposit with high mechanical integrity and with high productivity requires many physical and chemical occurrences in the welding arc. These occurrences are achieved in flux-related arc welding processes by the careful formulation of the welding flux. The welding flux must: • • • • • • • • •

STABILIZE ARC AND CONTROL ARC RESISTIVITY PROVIDE SLAG WITH PROPER MELTING TEMPERATURE PROVIDE LOW-DENSITY SLAG PERMIT USE OF DIFFERENT TYPES OF CURRENT AND POLARITY ADD ALLOYING ELEMENTS REFINE THE WELD POOL (DEOXIDATION AND DESULFURIZATION) PROVIDE PROPER VISCOSITY FOR OUT-OF-POSITION WELDING PROMOTE SLAG DETACHABILITY PRODUCE SMOOTH WELD CONTOUR

•

REDUCE SPATTER AND FUME

It is apparent, after considering the large number of necessary requirements, that a welding flux must be carefully and deliberately formulated to achieve an optimized performance. A number of different arc welding processes depend on welding fluxes. Each of these processes requires a different formulation. SMAW Fluxes. The typical constituents and their functions in electrode coatings for the SMAW process are given in

Table 5. It should be noted that the flux ingredients are based on additions of refined minerals of the earth. Natural minerals offer an economical method of keeping welding consumables at a reasonable cost. Table 6 presents the elemental content of these minerals. Table 7 gives the typical chemical compositions for flux coatings for three different SMAW electrodes.

TABLE 5 TYPICAL FUNCTIONS AND COMPOSITIONS OF CONSTITUENTS FOR SELECTED MILD STEEL SMAW ELECTRODE COATINGS

COATING CONSTITUENT

FUNCTION OF CONSTITUENT PRIMARY SECONDARY

CELLULOSE

SHIELDING GAS SHIELDING GAS SLAG FORMER SHIELDING GAS SLAG FORMER ARC STABILIZER SLAG FORMER EXTRUSION EXTRUSION

CALCIUM CARBONATE FLUORSPAR DOLOMITE TITANIUM DIOXIDE (RUTILE) POTASSIUM TITANATE FELDSPAR MICA CLAY SILICA ASBESTOS MANGANESE OXIDE IRON OXIDE IRON POWDER FERROSILICON FERROMANGANESE SODIUM SILICATE

SLAG FORMER SLAG FORMER SLAG FORMER SLAG FORMER DEPOSITION RATE DEOXIDIZER ALLOYING BINDER

... FLUXING AGENT FLUXING AGENT FLUXING AGENT ARC STABILIZER SLAG FORMER STABILIZER

COMPOSITION RANGE OF COATING ON ELECTRODE E6010, E6013 E7018 E7024 E7028 E6011 25-40 2-12 ... 1-5 ... ...

0-5

15-30

0-5

0-5

...

...

15-30

...

5-10

...

...

...

...

5-10

10-20

30-55

0-5

20-35

10-20

(A)

(A)

0-5

...

0-5

...

0-20

0-5

...

0-5

STABILIZER SLAG FORMER ...

... ...

0-15 0-10

... ...

0-15 ...

... ...

...

...

...

...

...

EXTRUSION

10-20

...

...

...

...

ALLOYING

...

...

...

...

...

...

...

...

...

...

CONTACT WELDING ... DEOXIDIZER FLUXING

...

...

25-40

40-55

40-55

... 5-10 20-30

... 5-10 5-10

5-10 2-6 0-5

0-5 5-10 0-10

2-6 2-6 0-5

POTASSIUM SILICATE

ARC STABILIZER

AGENT BINDER

(B)

5-15(B) 5-10

0-10

0-5

(A) REPLACES TITANIUM DIOXIDE (RUTILE) TO PERMIT USE WITH ALTERNATING CURRENT. (B) REPLACES SODIUM SILICATE TO PERMIT USE WITH ALTERNATING CURRENT TABLE 6 TYPICAL COMPOSITION OF COMMON MINERALS USED IN SMAW ELECTRODE COATINGS

MINERAL ILMENITE TALC BENTONITE SILICA, QUARTZ CELLULOSE ALUMINA MUSCOVITE, MICA(A) ACTINOLITE MAGNETITE HEMATITE RUTILE, TITANIA DOLOMITE FLUORSPAR, FLUORITE CRYOLITE LIME LIMESTONE, CALCITE, MARBLE ZIRCONIA FELDSPAR(A) CLAY(A) SODIUM SILICATE POTASSIUM SILICATE CHROMIC OXIDE

CHEMICAL COMPOSITION FEO · TIO2 3MGO · 4SIO2 · H2O COMPLEX AL, MG, CA, FE HYDROXIDES SIO2 (C6H10O5)X AL2O3 K2O · 3AL2 · 6SIO2 · 2H2O CAO · MGO · 2FEO · 4SIO2 FE3O4 FE2O3 TIO2 MGO · CAO · (CO2)2 CAF2 NA3ALF4 CAO CACO3 ZRO2 K2O · AL2O3 · 6SIO2 AL2O3 · 2SIO2 · 2H2O SIO2/NA2O (RATIO-3.22) SIO2 K2O (RATIO-2.11) CR2O3

Source: Ref 1

(A) ALTHOUGH THESE SUBSTANCES CAN HAVE SEVERAL CHEMICAL COMPOSITIONS, ONLY TYPICAL COMPOSITION IS GIVEN.

TABLE 7 CHEMICAL COMPOSITION OF COVERINGS USED IN ELECTRODES FOR SMAW WELDING OF MILD STEELS AND LOW-ALLOY STEELS

COMPOSITION, WT%(A)

ELECTRODE DESCRIPTION

AWS DESIGNATION

HIGH-CELLULOSE, GAS E6010 SHIELDED HIGH-TITANIUM GAS-SLAG E6012 SHIELD LOW-HYDROGEN IRON E7018 POWDER Source: Ref 1

(A)

AFTER BAKING.

CAO

TIO2

...

10.1 . . .

...

46.0 . . .

14.4 . . .

CAF2

SIO2

AL2O3

MGO

NA3ALF3

N3O

FEO

47.0 . . .

3.2

...

5.1 1.3

23.6 5.0

2.0

...

11.0 20.5 2.0

1.0

5.0

SI

MN

FE

CO + CO2

VOLATILE MATTER

MOISTURE

1.5 2.8 . . .

...

25.0

4.0

2.4 7.0

1.5 2.5 . . .

...

5.0

2.0

1.2 . . .

2.5 1.8 28.5 12.0

...

0.1

Flux-cored arc welding (FCAW) uses a hollow wire filled with flux reagents and ferro-additions. The two types of

flux-cored electrodes are carbon-dioxide-shielded flux-cored electrodes and self-shielded flux-cored electrodes (Ref 26). Table 8 gives typical compositions for the three types of carbon-dioxide-shielded cored electrodes. Table 9 gives the typical compositions for the four types of self-shielded flux-cored electrodes.

TABLE 8 TYPICAL FLUX COMPOSITIONS OF AVAILABLE CO2 SHIELDED FCAW ELECTRODES

ELECTRODE DESCRIPTION FLUX TYPE NUMB TYPE ER 1 TITANIA (NONBAS IC) 2 LIMETITANIA (BASIC OR NEUTRA L) 3 LIME (BASIC)

COMPOSITION, WT% AWS DESIGNAT ION

SIO2

E70T-1, E70T-2

AL2 O3

TIO 2

ZR O2

CA O

21. 0

2.1

40. 5

. . . 0. 7

E70T-1

17. 8

4.3

9.8

E70T-1, E70T-5

7.5

0.5

...

NA2 O

K2 O

CO2 (AS CARBONA TE)

C

1.6 1. 4

0.5

0. 6

20. 1

15. 8

...

6.2 9. 7

1.9 1. 5

...

0. 3

24. 7

13. 0

18. 0

. . . 3. 2

. . . 0. 5

2.5

1. 1

55. 0

7.2

20. 5

FE

MN

CAF 2

Source: Ref 26

TABLE 9 TYPICAL FLUX COMPOSITIONS OF AVAILABLE SELF-SHIELDED FCAW ELECTRODES

ELECTRODE DESCRIPTION TYPE FLUX NUMB TYPE ER 1 FLUORSP ARALUMIN UM 2 FLUORSP ARTITANIA 3 FLUORSP AR-LIMETITANIA 4 FLUORSP AR-LIME

COMPOSITION, WT% AWS DESIGNAT ION

SI O2

AL

E70T-4, E70T-7, E60T-8

0. 5

15. . . . 4

. . . . . 12. 0. . 6 4

E70T-3

3. 6

1.9 . . .

20. . . 4.5 0. 6 . 6

E70T-6

4. 2

1.4 . . .

E70T-5

6. 9

. . . 0.6

AL2 O3

CO2 (AS CARBON ATE)

FE

0.2 1. 2

0.4

0.4 3. 0

. . 63. . 5

0.1 0. 6

0.6

50. 4. 0 5

. . 22. . 0

14. 4. 7 0

2.2 . . . . . 0. . 6

2.1

50. 2. 5 0

2. 4

1.2 3. 2

. . . . . 0.6 0. . 3

1.3

58. 7. 0 9

. . 22. . 0

TIO 2

CA O

MG O

K2 O

NA2 O

C

M N

NI

CA F2

15. 3

Source: Ref 26 Submerged Arc Welding (SAW). In the SAW process, the flux drops from a hopper onto the work such that the welding arc is submerged beneath the granular flux, producing an arc cavity that contains metal vapors, silicon monoxide, manganese oxide, gaseous fluorides, and other higher-vapor-pressure components of the flux. This arc welding process has been recognized as one that produces very little, if any, fume. Submerged arc welding is limited mainly to the flat or

horizontal position, and requires significant setup time. It is very successful in manufacturing numerous similar parts, such as producing welded steel pipe. Submerged arc fluxes are made in three different forms: • • •

BONDED FLUXES MIX NONMETALLIC AND FERRO-ADDITIONS WITH A LOWTEMPERATURE BINDER INTO MIXTURES OF SMALL PARTICLES. AGGLOMERATED FLUXES ARE SIMILAR TO THE BONDED FLUXES EXCEPT THAT A CERAMIC GLASS BINDER, CURED AT HIGH TEMPERATURE, IS USED. FUSED FLUXES ARE MADE BY POURING A HOMOGENEOUS GLASS MIXTURE OF THE PROPER FLUX COMPOSITION INTO WATER, RESULTING IN A FRIT.

Generally speaking, changing from bonded to agglomerated to fused fluxes improves control of the weld metal composition, especially with respect to such impurities as hydrogen and oxygen. There are seven types of submerged arc fluxes (Table 10). Flux classification according to basicity is the result of observed correlations between weld metal oxygen concentration and flux composition. Table 11 gives some typical compositions for SAW fluxes.

TABLE 10 FLUXES USED FOR SAW APPLICATIONS

FLUX TYPE MANGANESE SILICATE

CONSTITUENTS MNO + SIO2 > 50%

BASICITY FLUX FORM ACID FUSED

ADVANTAGES MODERATE STRENGTH; TOLERANT TO RUST; FAST WELDING SPEEDS; HIGH HEAT INPUT; GOOD STORAGE

CALCIUMHIGH SILICA

CAO + MGO + SIO2 > 60%

ACID

AGGLOMERATED, FUSED

HIGH WELDING CURRENT; TOLERANT TO RUST

CALCIUM SILICATENEUTRAL

CAO + MGO + SIO2 > 60%

NEUTRAL

AGGLOMERATED, FUSED

CALCIUM SILICATELOW SILICA

CAO + MGO + SIO2 > 60%

BASIC

AGGLOMERATED, FUSED

ALUMINATE

AL2O3 + CAO +

BASIC

AGGLOMERATED

MODERATE STRENGTH AND TOUGHNESS; ALL CURRENT TYPES; TOLERANT TO RUST; SINGLE- OR MULTIPLE-PASS WELD GOOD TOUGHNESS WITH MEDIUM STRENGTH; FAST WELDING SPEEDS; LESS CHANGE IN COMPOSITION AND LOWER OXYGEN GOOD STRENGTH

LIMITATIONS LIMITED USE FOR MULTIPASS WELDING; USE WHERE NO TOUGHNESS REQUIREMENT; HIGH WELD METAL OXYGEN; INCREASE IN SILICON ON WELDING; LOW IN CARBON POOR WELD TOUGHNESS; USE WHERE NO TOUGHNESS REQUIREMENT; HIGH WELD METAL OXYGEN ...

COMMENTS ASSOCIATED MANGANESE GAIN; MAXIMUM CURRENT, 1100 A; HIGHER WELDING SPEEDS

NOT TOLERANT TO RUST; NOT USED FOR MULTIWIRE WELDING

...

NOT TOLERANT TO

USUALLY

DIFFER IN SILICON GAIN; SOME CAPABLE OF 2500 A; WIRES WITH HIGH MANGANESE ...

BASIC

MGO > 45%; AL2O3 > 20%

ALUMINA

BAUXITE BASE

NEUTRAL

AGGLOMERATED, FUSED

BASIC FLUORIDE

CAO + MGO + MNO + CAF2 > 50%; SIO2 22%; CAF2 15%

BASIC

AGGLOMERATED, FUSED

AND TOUGHNESS IN MULTIPASS WELDS; NO CHANGE IN CARBON; LOSS OF SULFUR AND SILICON LESS CHANGE IN WELD COMPOSITION AND LOWER OXYGEN THAN FOR ACID TYPE; MODERATE TO FAST WELDING SPEEDS VERY LOW OXYGEN; MODERATE TO GOOD LOWTEMPERATURE TOUGHNESS

RUST, LIMITED TO DC ELECTRODE POSITIVE

...

MAY PRESENT PROBLEMS OF SLAG DETACHABILITY; MAY PRESENT PROBLEM OF MOISTURE PICKUP

MANGANESE GAIN; MAXIMUM CURRENT 1200 A; GOOD MECHANICAL PROPERTIES ...

CAN BE USED WITH ALL WIRES, PREFERABLE DC WELDING; VERY GOOD WELD PROPERTIES

TABLE 11 TYPICAL COMPOSITIONS OF SMAW FLUXES

FLUX COMPOSITION, WT% AL2O3 SIO2 TIO2 MGO A 49.9 13.7 10.1 2.9 B 24.9 18.4 0.2 28.9 C 19.3 16.3 0.8 27.2 D 18.1 13.2 0.5 28.2 E 17.0 12.2 0.7 36.8

CAF2 5.7 24.2 23.6 31.8 29.2

CAO ... ... 9.8 4.5 0.7

MNO 15.1 1.8 0.08 0.1 8.9

NA2O 1.6 2.1 0.9 0.9 1.6

K2O 0.2 0.07 1.1 0.9 0.1

BASICITY INDEX (BI) 0.4 1.8 2.4 3.0 3.5

References cited in this section

1. G.E. LINNERT, CHAPTER 8, WELDING METALLURGY, VOL 1, AWS, 1965, P 367-396 26. FUMES AND GASES IN THE WELDING ENVIRONMENT, AWS, 1979 Nature and Behavior of Fluxes Used for Welding D.L. Olson, S. Liu, R.H. Frost, G.R. Edwards, and D.A. Fleming, Colorado School of Mines

References

1. G.E. LINNERT, CHAPTER 8, WELDING METALLURGY, VOL 1, AWS, 1965, P 367-396 2. C.E. JACKSON, FLUXES AND SLAGS IN WELDING, WELD. RES. BULL., NO. 190, 1973 3. T. LAU, G.C. WEATHERLY, AND A. MCLEAN, THE SOURCES OF OXYGEN AND NITROGEN CONTAMINATION IN SUBMERGED ARC WELDING USING CAO-AL2O3 BASED FLUXES, WELD. J., VOL 64 (NO. 12), 1985, P 343S-347S 4. T.H. NORTH, H.B. BELL, A. NOWICKI, AND I. CRAIG, SLAG/METAL INTERACTION, OXYGEN, AND TOUGHNESS IN SUBMERGED ARC WELDING, WELD. J., VOL 57 (NO. 3), 1978, P 63S-75S 5. N. CHRISTENSEN AND J. CHIPMAN, SLAG-METAL INTERACTION IN ARC WELDING, WELD. RES. BULL., NO. 15, JAN 1953, P 1-14 6. R.H. FROST, D.L. OLSON, AND S. LIU, PYROCHEMICAL EVALUATION OF WELD METAL INCLUSION EVOLUTION, PROC. 3RD INT. CONF. TRENDS IN WELDING, ASM INTERNATIONAL, JUNE 1992 7. C.A. NATALIE, D.L. OLSON, AND M. BLANDER, PHYSICAL AND CHEMICAL BEHAVIOR OF WELDING FLUXES, ANN. REV. MATER. SCI., VOL 16, 1986, P 389-413 8. J.E. INDACOCHEA, M. BLANDER, N. CHRISTENSEN, AND D.L. OLSON, CHEMICAL REACTIONS WITH FEO-MNO-SIO2 FLUXES, METALL. TRANS. B, VOL 16, 1985, P 237-245 9. U. MITRA AND T.W. EAGAR, SLAG-METAL REACTIONS DURING WELDING, METALL. TRANS. B, VOL 22, 1991, P 65-100 10. C.S. CHAI AND T.W. EAGAR, SLAG-METAL EQUILIBRIUM DURING SUBMERGED ARC WELDING,METALL. TRANS. B, VOL 12, 1981, P 539-547 11. N. CHRISTENSEN, METALLURGICAL ASPECTS OF ARC WELDING, WELD. J., VOL 27, 1949, P 373S-380S 12. O. GRONG, D.L. OLSON, AND N. CHRISTENSEN, CARBON OXIDATION IN HYPERBARIC MMA WELDING, MET. CONSTRUCT., VOL 17. DEC 1985, P 810R-814R 13. U. MITRA AND T.W. EAGAR, SLAG-METAL REACTIONS DURING SUBMERGED ARC WELDING OF ALLOY STEELS, METALL. TRANS. A, VOL 15, 1984, P 217-227

14. T.W. EAGAR, SOURCES OF WELD METAL OXYGEN CONTAMINATION DURING SUBMERGED ARC WELDING, WELD. J.,VOL 57, 1978, P 76S-80S 15. N. CHRISTENSEN, WELDING METALLURGY, LECTURE NOTES, NTH, 1979 16. O. GRONG, T.A. SIEWERT, T.A. MARTINS, AND D.L. OLSON, A MODEL FOR THE SILICONMANGANESE DEOXIDATION OF STEEL WELD METAL, METALL. TRANS. A, VOL 17 (NO. 10), 1985, P 1797-1807 17. T.H. NORTH, THE DISTRIBUTION OF MANGANESE BETWEEN SLAG AND METAL DURING SUBMERGED ARC WELDING, WELD. RES. ABROAD, VOL 23 (NO. 1), 1977, P 2-40 18. T. BONISZEWSKI, BASIC FLUXES AND DEOXIDATION IN SUBMERGED ARC WELDING OF STEEL, METAL. CONSTR. BRIT. WELD. J., VOL 6, 1974, P 128 19. L. DAVIS, AN INTRODUCTION TO WELDING FLUXES FOR MILD AND LOW ALLOY STEELS, WELDING INSTITUTE, 1981 20. MAKING, SHAPING AND TREATING OF STEEL, UNITED STATES STEEL CORPORATION, 1984 21. P.S. DUNN, C.A. NATALIE, AND D.L. OLSON, SOL-GEL FLUXES FOR FLUX CORED WELDING CONSUMABLES, J. MATER. ENERGY SYSTEMS, VOL 8 (NO. 2), 1986, P 176-184 22. L.I. SOROKIN AND Z.A. SIDLIN, THE EFFECT OF ALLOYING ELEMENTS AND OF MARBLE IN AN ELECTRODE COATING ON THE SUSCEPTIBILITY OF A DEPOSITED NICKEL CHROME METAL TO PORE FORMATION, SVAR. PROIZVOD., NO. 11, 1974, P 7-9 23. S.S. TULIANI, T. BONISZEWSKI, AND N.F. EATON, NOTCH TOUGHNESS OF COMMERCIAL SUBMERGED ARC WELD METAL, WELD. MET. FABR., VOL 37 (NO. 8), 1969, P 27 24. T.G.F. GRAY, J. SPENCE, AND T.H. NORTH, RATIONAL WELDING DESIGN, BUTTERWORTHS, 1975 25. D.L. OLSON, G.R. EDWARDS, AND S.K. MARYA, THE PHYSICAL AND CHEMICAL BEHAVIOR ASSOCIATED WITH SLAG DETACHABILITY DURING WELDING, FERROUS ALLOY WELDMENTS, VOL 67-70, TRANS TECH, 1992, P 253-268 26. FUMES AND GASES IN THE WELDING ENVIRONMENT, AWS, 1979

Shielding Gases for Welding Kevin A. Lyttle, Praxair, Inc.

Introduction THE SHIELDING GAS used in a welding process has a significant influence on the overall performance of the welding system. Its primary function is to protect the molten metal from atmospheric nitrogen and oxygen as the weld pool is being formed. The shielding gas also promotes a stable arc and uniform metal transfer. In gas-metal arc welding (GMAW) and flux-cored arc welding (FCAW), the gas used has a substantial influence on the form of metal transfer during welding. This, in turn, affects the efficiency, quality, and overall operator acceptance of the welding operation. The shielding gas interacts with the base material and with the filler material, if any, to produce the basic strength, toughness, and corrosion resistance of the weld. It can also affect the weld bead shape and the penetration pattern. Understanding the basic properties of a shielding gas will aid in the selection of the right shielding gas or gases for a welding application. Use of the best gas blend will improve the quality and may reduce the overall cost of the welding operation as well. Shielding Gases for Welding Kevin A. Lyttle, Praxair, Inc.

Basic Properties of a Shielding Gas The "controlled electrical discharge" known as the welding arc is formed and sustained by the establishment of a conductive medium called the arc plasma. This plasma consists of ionized gas, molten metals, slags, vapors, and gaseous atoms and molecules. The formation and structure of the arc plasma is dependent on the properties of the shielding gases used for welding. Table 1 lists the basic properties of gases used for welding (Ref 1).

TABLE 1 PROPERTIES OF SHIELDING GASES USED FOR WELDING

GAS

ARGON CARBON DIOXIDE HELIUM HYDROGEN NITROGEN OXYGEN

CHEMICAL MOLECULAR SPECIFIC DENSITY (A) SYMBOL WEIGHT GRAVITY g/ft3 g/l AR 39.95 1.38 0.1114 1.784 CO2 44.01 1.53 0.1235 1.978

IONIZATION POTENTIAL aj(b) eV 2.52 15.7 2.26 14.4

HE H2 N2 O2

3.92 2.16 2.32 2.11

4.00 2.016 28.01 32.00

0.1368 0.0695 0.967 1.105

0.0111 0.0056 0.782 0.0892

0.178 0.090 12.5 1.43

24.5 13.5 14.5 13.2

Source: Ref 1

(A) (B)

AT 100 KPA (1 ATM) AND 0 °C (32 °F); AIR = 1. 10-18 J.

The ionization potential is the energy, expressed in electron volts, necessary to remove an electron from a gas atom--

making it an ion, or an electrically charged gas atom. All other factors held constant, the value of the ionization potential decreases as the molecular weight of the gas increases. Arc starting and arc stability are greatly influenced by the ionization potentials of the component shielding gases used in welding process. A gas with a low ionization potential,

such as argon, can atoms into ions easily. Helium, with its significantly higher ionization potential, produces a harder to start, less stable arc. Although other factors are involved in sustaining the plasma, the respective energy levels required to ionize these gases must be maintained; as a consequence, the arc voltage is directly influenced. For equivalent arc lengths and welding currents, the voltage obtained with helium is appreciably higher than is with argon. This translates into more available heat input to the base material with helium than with argon. The thermal conductivity of a gas is a measure of how well it is able to conduct heat. It influences the radial heat loss

from the center to the periphery of the arc column as well as heat transfer between the plasma and the liquid metal. Argon, which has a low thermal conductivity, produces an arc that has two zones: a narrow hot core and a considerably cooler outer zone. The penetration profile of the weld fusion area then exhibits a narrow "finger" at the root and a wider top. A gas with a high thermal conductivity conducts heat outward from the core; this results in a wider, hotter arc core. This type of heat distribution occurs with helium, argon-hydrogen, and argon-carbon dioxide blends; it gives a more even distribution of heat to the work surface and produces a wider fusion area. Dissociation and Recombination. Shielding gases such as carbon dioxide, hydrogen, and oxygen are multiatom

molecules. When heated to high temperatures within the arc plasma, these gases break down, or dissociate, into their component atoms. They are then at least partially ionized, producing free electrons and current flow. As the dissociated gas comes into contact with the relatively cool work surface, the atoms recombine and release heat at that point. This heat of recombination causes multiatomic gases to behave as if they have a higher thermal conductivity, similar to that of helium. Dissociation and recombination do not occur with gases, such as argon, that consist of a single atom. Thus, at the same arc temperature, the heat generated at the work surface can be considerably greater with gases such as carbon dioxide and hydrogen. Reactivity/Oxidation Potential. The oxidizing nature of the shielding gas affects both welding performance and the

properties of the resultant weld deposit. Argon and helium are completely nonreactive, or inert, and thus have no direct chemical affect on the weld metal. Oxidizing or active gases, such as CO2 and oxygen, will react with elements in the filler metal or baseplate and will form a slag on the surface of the weld deposit. The loss of elements, such as manganese and silicon, from steel can affect the quality and cost of the weldment produced. Both weld strength and toughness generally decline as the oxidizing nature of the shielding gas increases. Additions of reactive gases such as oxygen or carbon dioxide enhance the stability of the arc and affect the type of metal transfer obtained. Metal droplet size is decreased, and the number of droplets transferred per unit time increases as the level of oxygen in the shielding gas increases. Oxygen reduces the molten weld bead surface tension, promoting better bead wetting and higher welding travel speeds. Small additions of CO2 work in a similar manner. The surface tension between the molten metal and its surrounding atmosphere has a pronounced influence on bead

shape. If the surface energy is high, a convex, irregular bead will result. Low values promote flatter beads with minimal susceptibility for undercutting. Pure argon is generally associated with high interfacial energy, producing a sluggish weld puddle and high, crowned bead. The addition of a small amount of a reactive gas, such as oxygen, lowers this surface tension and promote fluidity and better wetting of the base material; it does this without creating excessive oxidation of the weld metal. Gas Purity. Some metals, such as carbon steel and copper, have a relatively high tolerance for contaminants in the

shielding gas; others, such as aluminum and magnesium, are fairly sensitive to particular contaminants. Still others, such as titanium and zirconium, have an extremely low tolerance for any foreign constituent in the shielding gas. Depending on the metal being welded and the welding process used, very small quantities of gas impurities can significantly affect welding speed, weld surface appearance, weld bead solidification, and porosity levels. The effects of any given impurity are wide ranging, but weld quality and eventual fitness for purpose are major areas of concern. There is always a possibility that the gas, as delivered, is contaminated; however, it is far more likely that impurities will enters somewhere between the supply and the end-use points. For this reason, property designed piping systems and highquality hose are recommended for use with welding shielding gases. Typical industry minimum purity levels for welding gases are listed in Table 2 (Ref 2).

TABLE 2 TYPICAL GASES PURITY AND MOISTURE CONTENT OF SHIELDING GAS

ARGON CARBON DIOXIDE HELIUM HYDROGEN NITROGEN OXYGEN

PRODUCT STATE

GAS LIQUID GAS LIQUID GAS LIQUID GAS LIQUID GAS LIQUID INDUSTRIAL LIQUID

MINIMUM PURITY, %

99.995 99.997 99.5 99.8 99.95 99.995 99.95 99.995 99.7 99.997 99.5 99.5

MAXIMUM MOISTURE, PPM(A)

10 6 19 50 32 3 8 5 32 5 50 6

APPROXIMATE DEW POINT AT MAXIMUM MOISTURE CONTENT

°C -60 -64 -51 -58 -51 -69 -63 -65 -51 -65 -48 -64

°F -77 -83 -60 -73 -61 -92 -80 -86 -61 -86 -54 -83

Source: Ref 2

(A) MOISTURE SPECIFICATIONS ARE MEASURED AT FULL CYLINDER PRESSURE, THE PRESSURE AT WHICH THE CYLINDER IS ANALYZED. Gas density is the weight of the gas per unit volume. Density is one of the chief factors that influence shielding gas effectiveness. Basically, gases heavier than air, such as argon and carbon dioxide, require lower flow rates in use than do the lighter gases, such as helium, to ensure adequate protection of the weld puddle.

References cited in this section

1. N.E. LARSON AND W.F. MEREDITH, SHIELDING GAS SELECTION MANUAL, UNION CARBIDE INDUSTRIAL GASES TECHNOLOGY CORP., 1990, P 10 2. N.E. LARSON AND W.F. MEREDITH, SHIELDING GAS SELECTION MANUAL, UNION CARBIDE INDUSTRIAL GASES TECHNOLOGY CORP., 1990, P 11

Shielding Gases for Welding Kevin A. Lyttle, Praxair, Inc.

Characteristics of the Components of a Shielding Gas Blend To obtain a shielding gas that is suited to a specific application, a mix of gases is generally needed. Each basic gas contributes certain characteristics to the performance of the overall mix. Some gas blends have relatively specific areas of application and limited operating ranges; others can be used on many materials under a variety of welding conditions. Each component of the blend brings with it properties that are supplemented by the others to produce an enhanced level of performance.

Argon is inert or unreactive with respect to the materials present in the welding electrode. With its low ionization

potential, argon promotes easy arc starting and stable arc operation. Its lower thermal conductivity promotes the development of axial "spray" transfer in certain forms of GMAW. It is also used in applications where base material distortion must be controlled or where good gap-bridging ability is required. Helium. Unlike argon, helium is lighter than air and has a low density. Like argon, it is chemically inert and does not

react with other elements or compounds. Because of its high thermal conductivity and high ionization potential, more heat is transferred to the base material, thus enhancing the penetration characteristics of the arc. In many applications, it also allows higher weld travel speeds to be obtained. Because of its higher cost, helium is frequently combined with argon or argon mixtures to enhance the overall performance of the blend while minimizing its cost. Oxygen combines with almost all known elements except rare and inert gases; it vigorously support combustion. Small

amounts of oxygen are added to some inert mixtures to improve the stability of the welding arc developed as well as to increase the fluidity of the weld puddle. In the spray-transfer mode of GMAW, small additions of oxygen enhance the range over which this spatterless form of welding can be performed. The droplet size decreases and the number of drops transferred per unit time increases as oxygen is added to the blend. Carbon dioxide is a reactive gas that is commonly used alone in certain types of GMAW. Oxidation of the base material

and any filler electrode occurs readily. Carbon dioxide is added to argon blends to improve are stability, enhance penetration, and improve weld puddle flow characteristics. The higher thermal conductivity of carbon dioxide (because of the dissociation and recombination of its component parts) transfer more heat to the base material than does argon alone. A broader penetration pattern versus argon is obtained; however, base material distortion and lack of gap-bridging ability are possible problems. Hydrogen is the lightest known element and is a flammable gas. Explosive mixtures can be formed when certain

concentration of hydrogen are mixed with oxygen or air. It is added to inert gases to increase the heat input to the base material or for operations involving cutting and gouging. Because some materials are especially sensitive to hydrogenrelated contamination, its use is generally limited to special applications, such as the joining of stainless steels, and to plasma are cutting and gouging. Nitrogen is generally considered to be inert except at high temperatures. At arc welding temperatures, it will react with

some metals (e.g., aluminum, magnesium, steel, and titanium), so it is not used as a primary shielding gas. It can be used with other gases for some welding applications (e.g., copper) and is also widely used in plasma cutting. Shielding Gases for Welding Kevin A. Lyttle, Praxair, Inc.

Shielding Gas Selection In most welding applications, more than one shielding gas or gas blend can be used successfully. For example, there is no one optimal gas blend for joining carbon steels, but a considerable array of mixes are available depending on the specific requirements of the application. For some processes, such as gas-tungsten arc welding (GTAW) and plasma arc welding (PAW), the choices may be somewhat limited by the nature of the electrodes used and the materials being welded. However, for applications involving GMAW, a multitude of blends can be selected from when carbon steels are being joined. Determination of the best blend depends on a number of specific job-related needs. Accuracy of Gas Blends The accuracy with which gases are blended is a function of the way in which they are supplied. If the source of the gas is a high-pressure cylinder, the following generally applies:

1. ±10% RELATIVE, MINOR COMPONENT (REF 3)

2. ±0.5% ABSOLUTE FOR CONCENTRATIONS UP CONCENTRATIONS BETWEEN 5 AND 50% (REF 4)

TO

5%;

±10%

RELATIVE

FOR

For example, the mix accuracy of an Ar-2O2 blend would be Ar/1.8-2.2O2 (method 1) or Ar/1.5-2.5O2 (method 2). A blend of Ar-25CO2 would yield Ar/22.5-27.5CO2 by their method of calculation. When gas cylinders are properly filled with the appropriate blend, the components of that mixture will not separate unless the temperature of the environment is reduced far below normal working temperatures. If the gases are supplied from a liquid source, such as a bulk tank, the accuracy of the blend is a function of the mixing equipment used, but most likely falls within the ±10% minor component range. Shielding Gases for GMAW By far, the largest number of gas blends have been developed for GMAW, especially for joining carbon steel. These can be roughly divided into four categories: pure gases, argon-oxygen mixes, argon/carbon dioxide mixes, and three-part gas blends composed of either argon, helium, oxygen, carbon dioxide, or hydrogen. Table 3 contains suggestions for shielding gas selection based on material type, thickness, and mode of metal transfer.

TABLE 3 RECOMMENDED SHIELDING GAS SELECTION FOR GMAW

MATERIAL CARBON STEEL

THICKNESS TRANSFER MODE MM IN. 0.125

SHORT CIRCUITING

AR-15CO2 AR-25CO2 CO2

GLOBULAR

AR-25CO2 CO2

CONVENTIONAL SPRAY ARC

AR-1O2 AR-2O2 AR-5CO2

ADVANTAGES AND LIMITATIONS GOOD PENETRATION AND DISTORTION CONTROL TO REDUCE POTENTIAL BURNTHROUGH HIGHER DEPOSITION RATES WITHOUT BURNTHROUGH: MINIMUM DISTORTION AND SPATTER; GOOD PUDDLE CONTROL FOR OUT-OF-POSITION WELDING HIGH WELDING SPEEDS, GOOD PENETRATION AND PUDDLE CONTROL; APPLICABLE LOT OUTOF-POSITION WELDS SUITABLE FOR HIGHCURRENT AND HIGHSPEED WELDING; DEEP PENETRATION AND FAST TRAVEL SPEEDS, BUT WITH GREATER BURNTHROUGH POTENTIAL GOOD ARC STABILITY; PRODUCES A MORE FLUID PUDDLE AS O2

PULSED SPRAY

ALLOY STEEL

ALL SIZES

STAINLESS ALL SIZES STEEL, COPPER, NICKEL, AND CU-NI ALLOYS

AR-8CO2 AR-10CO2 AR-15CO2 AR-CO2-O2 BLENDS ARGON-5CO2 AR-HE-CO2 BLENDS AR-CO2-O2 BLENDS

SHORT CIRCUITING

AR-8CO2 AR-15CO2 AR-CO2-O2 BLEND

SPRAY ARC (HIGH-CURRENT DENSITY AND ROTATIONAL)

AR-2O2 AR-5O2 AR-CO2-O2 BLENDS AR-HE-CO2 BLENDS

PULSED SPRAY

AR-5CO2 AR-8CO2 AR-2O2

SHORTCIRCUITING TRANSFER

AR-HE-CO2 BLENDS HE-AR-CO2 BLENDS AR-1O2 AR-2O2

INCREASES; GOOD COALESCENCE AND BEAD CONTOUR, GOOD WELD APPEARANCE AND PUDDLE CONTROL USED FOR BOTH GAGE AND OUT-OF-POSITION WELDMENTS; ACHIEVES GOOD PULSED SPRAY STABILITY OVER A WIDE RANGE OF ARC CHARACTERISTICS AND DEPOSITION RANGES HIGH WELDING SPEEDS; GOOD PENETRATION AND PUDDLE CONTROL; APPLICABLE FOR OUTOF-POSITION WELDS; SUITABLE FOR HIGHCURRENT AND HIGHSPEED WELDING REDUCES UNDERCUTTING; HIGHER DEPOSITION RATES AND IMPROVED BEAD WETTING; DEEP PENETRATION AND GOOD MECHANICAL PROPERTIES USED FOR BOTH LIGHTGAGE AND HEAVY OUT-OF-POSITION WELDMENTS; ACHIEVES GOOD PULSED SPRAY STABILITY OVER A WIDE RANGE OF ARC CHARACTERISTICS AND DEPOSITION RANGES LOW CO2 CONTENTS IN HELIUM MIX MINIMIZE CARBON PICKUP, WHICH CAN CAUSE INTERGRANULAR CORROSION WITH SOME ALLOYS; HELIUM IMPROVES WETTING ACTION; CO2 CONTENTS >5% SHOULD BE USED WITH CAUTION ON

ALUMINUM, TITANIUM, AND OTHER REACTIVE METALS

≤ 13

>13

≤

>

1 2

1 2

SPRAY ARC

AR-HE-CO2 BLENDS AR-1O2 AR-2O2

PULSED SPRAY

AR-HECO2BLENDS AR-1O2 AR-2O2

SPRAY ARC

ARGON

SPRAY ARC

75HE-25AR 50HE-50AR

SPRAY ARC

HELIUM 50HE-25AR

PULSED SPRAY

ARGON

SOME ALLOYS; APPLICABLE FOR ALL POSITION WELDING GOOD ARC STABILITY; PRODUCES A FLUID BUT CONTROLLABLE WELD PUDDLE; GOOD COALESCENCE AND BEAD CONTOUR; MINIMIZES UNDERCUTTING ON HEAVIER THICKNESSES USED FOR BOTH LIGHTGAGE AND HEAVY OUT-OF-POSITION WELDMENTS; ACHIEVES GOOD PULSED SPRAY STABILITY OVER A WIDE RANGE OF ARC CHARACTERISTICS AND DEPOSITION RANGES BEST METAL TRANSFER, ARC STABILITY, AND PLATE CLEANING; LITTLE OR NO SPATTER; REMOVES OXIDES WHEN USED WITH DCEP (REVERSE POLARITY) HIGH HEAT INPUT; PRODUCES FLUID PUDDLE, FLAT BEAD CONTOUR, AND DEEP PENETRATION; MINIMIZES POROSITY HIGH HEAT INPUT; GOOD FOR MECHANIZED WELDING AND OVERHEAD; APPLICABLE TO HEAVY SECTION WELDING GOOD WETTING; GOOD PUDDLE CONTROL

Argon. Pure argon is generally used on nonferrous base metal, such as aluminum, nickel, copper, and magnesium alloys,

and on reactive metals, such as titanium. Argon provides excellent arc stability, penetration, and bead profile when joining these materials. Its low ionization potential results in easy arc starting. Argon produces a constricted arc column with high current density, which concentrates the arc energy over a small area; deep, fingerlike penetration results. Carbon Dioxide. A reactive gas, carbon dioxide is generally used only for joining carbon steel. It is readily available and

relatively inexpensive. Because CO2 will not support spray transfer, deposition efficiency is lower and spatter and fume

levels are higher than with argon blends. Weld bead surfaces are more oxidized and irregular in shape. The higher ionization potential of CO2 and its characteristic dissociation upon heating provide greater weld fusion and penetration while still achieving acceptable mechanical properties. Helium. Because of its higher thermal conductivity, helium can provide additional heat input to the base material while

still maintaining an inert atmosphere. Wetting action, depth of fusion, and travel speed can be improved over comparable argon levels. This advantage is most frequently utilized in the welding of heavier sections of aluminum, magnesium, and copper alloys. Argon-Oxygen. The addition of a small amount of oxygen to a argon greatly stabilizes the welding arc, increases the

filler metal droplet rate, lowers the spray transition current, and influences bead shape. The weld pool is more fluid and stays molten longer, allowing the metal to flow out toward the edges of the weld (Fig. 1a).

FIG. 1 EFFECT OF SHIELDING GAS BLENDS ON WELD PROFILE USING DIRECT CURRENT ELECTRODE POSITIVE (DCEP). (A) ARGON VERSUS ARGON-OXYGEN. (B) CARBON DIOXIDE VERSUS ARGON/CARBON DIOXIDE. (C) HELIUM VERSUS ARGON-HELIUM. SOURCE: REF 5

The most common blends contain 1, 2, 5, or 8% O2 in argon. Increasing oxygen improves arc stability and makes higher travel speeds possible by enhanced puddle fluidity. Some increased alloy loss and a greater chance of undercut occur as the oxygen level is increased, especially beyond 5%. Argon/carbon dioxide blends are primarily used for carbon and low-alloy steels and have limited use for stainless

steels. The addition of CO2 to argon produces results similar to the addition of oxygen, but is also broadens the penetration pattern as the CO2 content is increased (Fig. 1b). Above a range of 18 to 20% CO2, spraylike transfer can no longer be obtained; short-circuiting/globular transfer with somewhat increased spatter levels is found from this point up to approximately 50% CO2 in argon. The most common blends for spray transfer are argon plus 5, 8, 10, or 13 to 18% CO2. With increased CO2 content, the more fluid weld puddle permits higher weld travel speeds. Mixes with higher carbon dioxide levels can also be used for short-circuiting transfer--commonly, argon plus 20 or 25% CO2. Mixtures in this range provide an optimum droplet frequency for minimum spatter when small-diameter (0.9 and 1.2 mm, or 0.035 and 0.045 in.) wire is used. Argon-Helium. Helium is often mixed with argon to obtain the advantages of both gases. These blends are primarily

used for nonferrous base materials, such as aluminum, copper, and nickel alloys. Helium increases the heat input to the base material and thus is used for joining thick, thermally conductive plates. As the helium percentage increases, the arc voltage, spatter, and weld width-to-depth ratio increase (Fig. 1c). The most common blends contain 25, 50, or 75% He in argon. The highest percentage of helium is used for joining thick (>50 mm, or 2 in.) plate material, especially aluminum and copper. Higher travel speeds can be obtained using heliumenhanced blends. Argon/Oxygen/Carbon Dioxide. Mixtures containing three gas components are versatile because of their ability to

operate in short-circuiting, globular, and spray-transfer modes. These blends are generally proprietary, and manufacturers' recommendations should be followed for their proper use. Blends of argon, carbon dioxide, and oxygen are generally used to join carbon and alloy steels.

Argon/Helium/Carbon Dioxide. Helium and carbon dioxide additions to argon increase the heat input to the base

metal, which improves wetting, puddle fluidity, and weld bead profile. With helium plus CO2 additions less than 40%, good spray transfer is obtained for carbon and low-alloy steel welding. Some increase in tolerance of base material surface contamination is also noted. When the helium content exceeds 50 to 60%, transfer is restricted to short-circuiting and globular. Blends in which the CO2 content is relatively low ( ≤ 5%) are generally used for the joining of stainless steels without any loss in corrosion resistance. Argon/Carbon Dioxide/Hydrogen. Three-part blends of this design are intended for the joining of austenitic stainless

in the spray or short-circuiting transfer modes. Because of the addition of hydrogen, these blends should not be used for carbon steel. The carbon dioxide and hydrogen increase the heat input to the base material and improve bead shape characteristics, as well as promote higher welding travel speeds. Shielding Gas Selection for FCAW Carbon Dioxide. The majority of large-diameter (>1.6 mm, or

1 in.) wires that use a shielding gas use carbon dioxide. 16

Some smaller diameter wires are formulated to operate in 100% CO2. The arcs are generally stable and provide a globular transfer over the usable operating range. Good performance over rust and mill scale on the plate surface is obtained with these large-diameter wires and CO2 shielding. Argon/Carbon Dioxide. A significant number of small-diameter ( ≤ 1.6 mm, or

1 in.) cored wires are shielded with the 16

blends of argon with 15 to 50% CO2. These blends provide better out-of-position weld puddle control versus that of CO2. To obtain the best performance from a particular cored wire, check the manufacturer's product literature for the recommended gas blend. Shielding Gas Selection for GTAW Argon. The most commonly used gas for GTAW, argon exhibits low thermal conductivity, which produces a narrow,

constricted arc column; this allows greater variations in arc length with minimal influence on arc power or weld bead shape. Its low ionization potential provides good arc starting characteristics and good arc stability using the direct current electrode negative (DCEN) power connection plus superior arc cleaning action and bead appearance when ac power is used. Argon is the most commonly selected gas for DCEN welding of most materials and ac manual welding of aluminum. Helium. The high thermal conductivity and ionization potential of helium make it suitable for the high-current joining of

heavy sections of heat-conductive materials such as aluminum. Helium increases the penetration of the weld as well as its width. It also allows the use of higher weld travel speeds. Argon-Helium. Blends of argon and helium are selected to increase the heat input to the base material while maintaining

favorable arc stability and superior arc starting characteristics. Blends of 25, 50, and 75% He in argon are commonly used. Argon-Hydrogen. Hydrogen is added to argon to enhance its thermal properties. The slightly reducing atmosphere

improves weld puddle wetting and reduces some surface oxides to produce a cleaner weld surface. To minimize problems associated with arc starting, additions of hydrogen are generally limited to 5 to 15%. These blends are primarily used to join some stainless steels, nickel, and nickel alloys. These mixtures should not be used to join alloy steels; delayed weld cracking may result. Argon/2-5H2 is used in manual welding applications on materials thicker than 1.6 mm (

1 in.). Additions of 10 to 15% 16

H2 are used in mechanized applications, such as those found in the manufacture of stainless steel tubing. Caution: Special safety precautions are required when mixing argon and hydrogen. Do not attempt to mix argon and hydrogen from separate cylinders. Always purchase ready-mixed hydrogen blends from a qualified supplier.

Shielding Gas Selection for PAW The physical configuration of the PAW system requires the use of two gases: a "plasma" or orifice gas and a shielding gas. The primary role of the plasma gas, which exits the torch through the center orifice, is to control arc characteristics and shield the electrode. The shielding gas, introduced around the periphery of the arc, shields or protects the weld area. In many applications, the shielding gas is also partially ionized to enhance the performance of the plasma gas. Low-Current (1.6 mm, or

1 in.) wires that use a shielding gas use carbon dioxide. 16

Some smaller diameter wires are formulated to operate in 100% CO2. The arcs are generally stable and provide a globular transfer over the usable operating range. Good performance over rust and mill scale on the plate surface is obtained with these large-diameter wires and CO2 shielding. Argon/Carbon Dioxide. A significant number of small-diameter ( ≤ 1.6 mm, or

1 in.) cored wires are shielded with the 16

blends of argon with 15 to 50% CO2. These blends provide better out-of-position weld puddle control versus that of CO2. To obtain the best performance from a particular cored wire, check the manufacturer's product literature for the recommended gas blend. Shielding Gas Selection for GTAW Argon. The most commonly used gas for GTAW, argon exhibits low thermal conductivity, which produces a narrow,

constricted arc column; this allows greater variations in arc length with minimal influence on arc power or weld bead shape. Its low ionization potential provides good arc starting characteristics and good arc stability using the direct current electrode negative (DCEN) power connection plus superior arc cleaning action and bead appearance when ac power is used. Argon is the most commonly selected gas for DCEN welding of most materials and ac manual welding of aluminum. Helium. The high thermal conductivity and ionization potential of helium make it suitable for the high-current joining of

heavy sections of heat-conductive materials such as aluminum. Helium increases the penetration of the weld as well as its width. It also allows the use of higher weld travel speeds. Argon-Helium. Blends of argon and helium are selected to increase the heat input to the base material while maintaining

favorable arc stability and superior arc starting characteristics. Blends of 25, 50, and 75% He in argon are commonly used. Argon-Hydrogen. Hydrogen is added to argon to enhance its thermal properties. The slightly reducing atmosphere

improves weld puddle wetting and reduces some surface oxides to produce a cleaner weld surface. To minimize problems associated with arc starting, additions of hydrogen are generally limited to 5 to 15%. These blends are primarily used to join some stainless steels, nickel, and nickel alloys. These mixtures should not be used to join alloy steels; delayed weld cracking may result. Argon/2-5H2 is used in manual welding applications on materials thicker than 1.6 mm (

1 in.). Additions of 10 to 15% 16

H2 are used in mechanized applications, such as those found in the manufacture of stainless steel tubing. Caution: Special safety precautions are required when mixing argon and hydrogen. Do not attempt to mix argon and hydrogen from separate cylinders. Always purchase ready-mixed hydrogen blends from a qualified supplier.

Shielding Gas Selection for PAW The physical configuration of the PAW system requires the use of two gases: a "plasma" or orifice gas and a shielding gas. The primary role of the plasma gas, which exits the torch through the center orifice, is to control arc characteristics and shield the electrode. The shielding gas, introduced around the periphery of the arc, shields or protects the weld area. In many applications, the shielding gas is also partially ionized to enhance the performance of the plasma gas. Low-Current ( γSL > γLV

(EQ 2)

For soldering, the vapor phase will be replaced in nearly all processes by flux (that is, γsv is replaced by γsf and γlv is replaced by γlf). For optimum wetting, the contact angle, θ, must be minimized and, therefore, γsv must be maximized. Oxides and contamination lower the value of γsv (or γsf). A function of the flux is to remove or disperse oxides and contaminants, thereby increasingly γsv. An important function of flux is to lower the value of γlf (or γlv). The flux also maintains a local environment around the surfaces being joined, protecting them from reoxidation during soldering, as well as enhancing heat transfer from the heating source to the substrate and the solder. Equation 1 shows that wetting is also enhanced by minimizing γlv, a value commonly referred to as the surface tension of the solder. Although γlv is a function of the solder composition, the flux covering the liquid solder, and the temperature, it is normally close to 0.4 J/m2 (4 × 10-5 Btu/ft2). A typical example of nonwetting occurs when a liquid metal droplet is placed on a nonmetallic surface, such as copper oxide. Because there are no metallic bonds at the oxide surface, the liquid metal will have little tendency to interact with that surface. The metal droplet will behave as though it is repelled, and will try to ball up to minimize the area of contact with the nonmetallic surface. One method for quantifying wettability is to observe the behavior of solder droplets on the metal surface. The contact angle formed between the surface of the liquid solder and the surface of the solid can be

measured (Fig. 5). If the solder droplet forms an angle, θ, of less than 90°, then the solder is said to wet the metal. If the angle is greater than 90°, the solder is nonwetting. This is one method used to qualitatively assess solderability. The formation of intermetallic compounds can affect Young's equation because it alters the value of γsl. The influence of this term has been recognized (Ref 6), but the magnitude has not yet been calculated quantitatively. Surfaces of a metal crystal have a higher free energy than the bulk of the crystal because of the existence of unsatisfied metallic bonds. When a liquid metal or alloy is brought into contact with this solid surface, the liquid will proceed to interact with the solid to satisfy those dangling surface bonds and thus reduce the surface free energy of the system. To the extent that a liquid solder can satisfy these surface bonds, it will wet the solid metal and spread across its surface. If no exchange of bonding energy takes place, then wetting will not occur and the solder will tend to ball up on the surface of the solid, minimizing the area of contact between liquid and solid. Information about selecting solders for specific applications is provided in the article "General Soldering" in this Volume.

References cited in this section

3. R.J. KLEIN WASSINK, SOLDERING IN ELECTRONICS, ELECTROCHEMICAL PUBLICATIONS LTD., 1989 6. A.D. ROMIG, JR., Y.A. CHANG, J.J. STEPHENS, ET AL., PHYSICAL METALLURGY OF SOLDERSUBSTRATE REACTIONS, SOLDER MECHANICS, THE MINERALS, METALS, & MATERIALS SOCIETY, 1991, P 30-32 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Guidelines for Flux Selection A flux promotes solder wetting of the base materials by: • • •

REMOVING TARNISH FILMS FROM PRECLEANED SURFACES PREVENTING OXIDATION DURING THE SOLDERING OPERATION LOWERING THE SURFACE TENSION OF THE SOLDER

Successful establishment of a solder joint requires that the liquid solder make contact with the metal to which it is to be joined, so that wetting can be initiated. Unfortunately, almost all of the metals involved in soldering are oxidized during exposure to elevated temperatures in air. This prevents metal-to-metal contact, as well as the wetting and formation of a metallurgical bond, unless the oxides are removed. Fluxes are used in soldering to remove such films and to protect the surfaces against reoxidation during soldering. In addition to cleaning metal surfaces, a flux also: • • •

AIDS IN THE WETTING PROCESS IMPROVES HEAT TRANSFER (ESPECIALLY IN SOLDER PASTES) CARRIES OXIDES AND SOLID DEBRIS AWAY FROM THE JOINT

Fluxes are chemical (liquid, solid, or gaseous materials) that remove oxide layers from the base metal and solder. When heated, fluxes either promote or accelerate the wetting of metals by solder. Flux selection usually depends on the ease with which a material can be soldered. Rosin fluxes are used with solderable base metals or with metals that are precoated

with a solderable base metal or with metals that are precoated with a solderable finish. Inorganic fluxes are often used on metals such as stainless steel. Table 1 indicates the relative ease with which a number of alloys and metals can be soldered, based on flux requirements.

TABLE 1 RELATIVE SOLDERABILITY OF SELECTED METALS AND ALLOYS AS A FUNCTION OF FLUX TYPE USED Base Metal, Alloy, Or Applied Finish

Flux Type ROSIN ORGANIC

INORGANIC

SPECIAL FLUX AND SOLDER

Aluminum Aluminum-Bronze Beryllium Beryllium-Copper Brass Cadmium Cast Iron Chromium Copper Copper-Chromium Copper-Nickel Copper-Silicon Gold Inconel Lead Magnesium Manganese-Bronze (High Tensile) Monel Nickel Nickel-Iron Nichrome Palladium Platinum Rhodium Silver Stainless Steel Steel Tin Tin-Bronze Tin-Lead Tin-Nickel Tin-Zinc Titanium Zinc Zinc Die Castings

... ... ... X X X ... ... X ... X ... X ... X ... ... ... ... ... ... X X ... X ... ... X X X ... X ... ... ...

... ... ... X X X ... ... X X X X X ... X ... ... X X X ... X X X X X X X X X X X ... X ...

X X ... ... ... ... X ... ... ... ... ... ... X ... ... ... ... ... ... X ... ... ... ... ... ... ... ... ... ... ... ... ... ...

... ... ... X X X ... ... X ... X ... X ... X ... ... X X X ... X X ... X X ... X X X X X ... X ...

Soldering Not Recommended(A)

... ... X ... ... ... ... X ... ... ... ... ... ... ... X X ... ... ... ... ... ... ... ... ... ... ... ... ... ... ... X ... X

(A) WITH PROPER PROCEDURES, SUCH AS PRECOATING, MOST METALS CAN BE SOLDERED Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Types of Fluxes Fluxes can be categorized by their chemical makeup (Ref 2, 3, 9, 10, 11): • • • •

ROSIN-BASE FLUXES ORGANIC FLUXES INORGANIC FLUXES SYNTHETICALLY ACTIVATED FLUXES (RESINS)

These materials vary in their activity (that is, aggressiveness). Rosin-base fluxes are very mild, whereas inorganic fluxes can be extremely active and corrosive. Because of the potential for corrosion by inorganic fluxes in electronic assemblies, the rosin-base fluxes (with their low flux activity) are favored by the electronics industry. Rosin Fluxes. Water-white rosin dissolved in a suitable organic solvent is the closest model of a flux with noncorrosive

residue. Rosin fluxes possess important physical and chemical properties that make them particularly suitable for use in the electrical industry. They are solid and inactive chemically at room temperature, but at soldering temperatures they gain sufficient activity to remove weakly adherent oxides from the noble metals gold and silver, as well as copper. Nonactivated Rosin Flux. Although its noncorrosive nature has led to its widespread use in microelectronics, rosin-

base flux is lacking in chemical activity, which essentially limits its use to precleaned parts and to only a few metals that do not have adherent oxides. The active constituent, abietic acid (C20H30O2), with a melting point of 173 °C (343 °F), becomes mildly active at soldering temperatures ranging from 177 to 316 °C (350 to 601 °F). The residue is hard, nonhygroscopic, electrically nonconductive, and noncorrosive. Mildly Activated Rosin Flux. The very low chemical activity of rosin-base flux can be countered by the addition of

activators, usually organic acids or amines. The end product, rosin mildly activated (RMA) flux, is still essentially noncorrosive, but is sufficiently active to remove oxides more reliably than the rosin-base flux. In addition, the residues are noncorrosive. However, care must be exercised in cases of high-reliability applications. It is advisable to specify that these fluxes be halide-free, in order to ensure that specific safety requirements are met in critical components. Even so, cleaning is now required to remove the flux residue. Both polar and nonpolar solvents have been used. These RMA fluxes are preferred for military, telephone, and other high-reliability electronic products. Activated rosin fluxes were developed to provide more chemically active fluxes for mass-produced electronics, such as

packaged components. Most mass-produced electronics are manufactured using RMA fluxes. The use of chlorides in these fluxes requires effective cleaning after soldering to prevent corrosion and electrical leakage, because the presence of chloride ions in flux residue makes it conductive. Because of the chemistry of the activators, a double solvent cleaning is normally required. Activated rosin fluxes are widely used in commercial electronics and in high-reliability applications where the residue can be completely removed after soldering. Inorganic Fluxes. Inorganic acids and salts that are highly corrosive and extremely active compose this class of fluxes.

In many instances, they are not acceptable for use on electronics, but are suitable for plumbing and industrial applications. Inorganic fluxes do enable the soldering of ferrous alloys and high-nickel alloys used in electronic packages and hermetic enclosures. The difficulty of chloride ion removal has led to the gradual abandonment of inorganic fluxes, even for tinning purposes. Inorganic fluxes are used to optimum advantage where conditions require rapid and highly active fluxing action. They can be applied as solutions, pastes, or dry salts in many general soldering applications. They function equally well with torch, oven, resistance, or induction soldering methods, because they neither char nor burn. These fluxes can be formulated to provide stability over the entire soldering temperature range. One distinct disadvantage of inorganic fluxes is that their residue remains chemically active after soldering. If this residue is not removed, then severe corrosion can occur at the joint. Adjoining areas can also be attacked by residues from the spray of flux and from flux vapors. Fluxes that contain ammonium salts can cause stress-corrosion cracking (SCC) when soldering brass, as well as most iron-nickel alloys.

Organic acid fluxes, although less active than inorganic fluxes, are effective at soldering temperatures ranging from 90

to 320 °C (195 to 610 °F). They consist of organic acids and bases and, often, hydrohalides. They are active at soldering temperatures, but the period of activity is short because of their susceptibility to thermal decomposition. Their tendency to volatilize, char, or burn when heated limits their use with torch or flame heating. When these fluxes are properly used, their residues are relatively inert and can be removed with water. Organic acid fluxes are particularly useful in applications where controlled quantities of flux can be applied and where sufficient heat can be used to fully decompose or volatilize the corrosive constituents. Caution is necessary to prevent undecomposed flux from spreading to insulating sleeving. Care must also be taken when soldering in closed systems where corrosive fumes can condense on critical parts of the assembly.

References cited in this section

2. H.H. MANKO, SOLDERS AND SOLDERING, 2ND ED., MCGRAW-HILL, 1964 3. R.J. KLEIN WASSINK, SOLDERING IN ELECTRONICS, ELECTROCHEMICAL PUBLICATIONS LTD., 1989 9. J.F. SHIPLEY, INFLUENCE OF FLUX, SUBSTRATE AND SOLDER COMPOSITION ON SOLDER WETTING, WELD. J., VOL 54 (NO. 10), OCT 1975, P 357S-362S 10. H.H. MANKO, SOLDERING FLUXES--PAST AND PRESENT, WELD. J., VOL 52 (NO. 3), MARCH 1973, P 163-166 11. R.W. WOODGATE, THE HANDBOOK OF MACHINE SOLDERING, 2ND ED., JOHN WILEY & SONS, 1988 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Flux Evaluation Several tests are used to evaluate the relative activity of fluxes. The most common are the copper mirror test, the halide test, a surface conductivity test, and a test based on the resistivity of a water extract present in the flux. In all cases, a measure of the chemical activity of the flux is obtained. Because the fluxes all depend on other properties (for example, viscosity, ease of removal, and thermal behavior), chemical activity is only one factor in flux selection for a given application. Information about selecting fluxes for specific applications is provided in the article "General Soldering" in this Volume. Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Joint Design Joints should be designed to fulfill the requirements of the finished assembly, as well as to permit the application of the flux and solder. Joint design should maintain proper clearance during heating and upon solidification of the filler metal. Special fixtures may be necessary or the units can be crimped, clinched, wrapped, or otherwise held together. The selection of a joint design for a specific application will primarily depend on the service requirements of the assembly. It may also depend on such factors as the heating method to be used, the fabrication techniques utilized prior to soldering, the number of items to be soldered, and the method used to apply the solder. In general, solders have low

strength when compared with the metals for which they are used to join. Therefore, the soldered joint should be designed to avoid dependence on solder strength. The necessary strength can be provided by shaping the parts to be joined so that they engage or interlock, requiring the solder only to bond, seal, and stiffen the assembly. Figure 6 shows joint designs commonly used for soldering applications. The lap joint and the lock seam joint are used when soldering sheets. Lap joints are also applied to join pipes. Actually, the lap type of joint should be used whenever possible because it offers the best chance to obtain joints with maximum strength. It should also be used whenever a seal is required. Butt joints should be avoided whenever possible.

FIG. 6 BASIC JOINT CONFIGURATIONS USED FOR SOLDERING APPLICATIONS. (A) LAP JOINT. (B) LOCK SEAM JOINT. (C) BUTT JOINT. (D) PIPE JOINT

An important factor in joint design is the manner in which the solder will be applied to the joint. The designer must consider the number of joints per assembly and the number of assemblies to be manufactured. For limited production using a manual soldering process, the solder can be face-fed into the joint with few problems. However, for a large production lot of assemblies containing multiple joints, an automated process such as condensation/vapor-phase reflow soldering can be advantageous (Ref 12, 13, 14). In this case, the design must provide for accessible joints that are suitable for automated fluxing, soldering, and cleaning.

There should be enough clearance between the parts being joined to allow the solder to be drawn into the space by capillary action, but not so much that the solder is unable to fill the gap. Joint clearances ranging from 0.075 to 0.150 mm (0.0030 to 0.059 in.) are preferred for optimum strength, but variations are allowed in specific cases. It is often necessary to fabricate sample parts and to test the joints to ensure their producibility and capability regarding strength properties (Ref 2, 3, 10, 15, 16, 17). Other commonly used soldered joints are shown in Fig. 7; self-jigging joints are shown in Fig. 8. The most definitive work available on the loading of soldered joints is that used for the load-carrying capabilities of copper tube with sleevetype joints or fittings. Conservative joint design requires that only 50% of the joint be considered filled. Under normal conditions, it would be considered good soldering practice if 70% or more of the joint consists of sound joint material (Ref 2, 3, 4, 5, 6, 11, 15, 18). Additional information about joint configurations is provided in the article "General Soldering" in this Volume.

FIG. 7 JOINT DESIGNS FREQUENTLY USED IN SOLDERING OPERATIONS. SOURCE: REF 5

FIG. 8 METHODS THAT CAN BE USED TO MAKE SOLDER JOINTS SELF-JIGGING. SOURCE: REF 5

References cited in this section

2. H.H. MANKO, SOLDERS AND SOLDERING, 2ND ED., MCGRAW-HILL, 1964 3. R.J. KLEIN WASSINK, SOLDERING IN ELECTRONICS, ELECTROCHEMICAL PUBLICATIONS LTD., 1989 4. L.P. LAMBERT, SOLDERING FOR ELECTRONIC ASSEMBLIES, MARCEL DEKKER, 1987 5. SOLDERING MANUAL, AWS, 1978 6. A.D. ROMIG, JR., Y.A. CHANG, J.J. STEPHENS, ET AL., PHYSICAL METALLURGY OF SOLDERSUBSTRATE REACTIONS, SOLDER MECHANICS, THE MINERALS, METALS, & MATERIALS SOCIETY, 1991, P 30-32 10. H.H. MANKO, SOLDERING FLUXES--PAST AND PRESENT, WELD. J., VOL 52 (NO. 3), MARCH 1973, P 163-166 11. R.W. WOODGATE, THE HANDBOOK OF MACHINE SOLDERING, 2ND ED., JOHN WILEY & SONS, 1988 12. T. THOMPSON, CONDENSATION/VAPOR-PHASE REFLOW SOLDERING, ASS. ENG., JUNE 1977, P 44-47 13. T.Y. CHU, A GENERAL REVIEW OF MASS SOLDERING METHODS, INS. CIRC., NOV 1976 14. R.C. PFAHL, JR., J.C. MOLLENDORF, AND T.Y. CHU, CONDENSATION SOLDERING, WELD. J., VOL 54 (NO. 1), 1975 15. E. LIEBERMAN, MODERN SOLDERING AND BRAZING TECHNIQUES, BUSINESS NEWS PUBLISHING CO., 1988 16. "SOLDER METAL SPECIFICATION," ASTM B 32, ANNUAL BOOK OF ASTM STANDARDS, PART B, ASTM 17. D.R. FREAR, W.B. JONES, AND K.R. KINSMAN, ED., SOLDER MECHANICS, TMS, 1991, P 30-41 18. PROC. IEEE 41ST ELECT. COMP. & TECH. CONF., IEEE, 1991 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Precleaning and Surface Preparation An unclean surface will not permit the solder to flow, which makes soldering either difficult or impossible and contributes to the formation of a poor joint. Materials such as oil, grease, paint, pencil markings, drawing, and cutting lubricants, general atmospheric dirt, oxide, and rust films must be removed before soldering. The importance of cleanliness cannot be overemphasized in ensuring sound soldered joints. Fluxing alone cannot substitute for adequate precleaning (Ref 19). Because cleaning methods are often designed for a specific soldering operation, their suitability for a particular application should be investigated thoroughly. Degreasing. Either solvent or alkaline degreasing is recommended for the cleaning of oily or greasy surfaces. Of the

solvent degreasing methods, the vapor condensation of the trichloroethylene-type solvents probably minimizes residual film on the surface. In the absence of vapor degreasing apparatus, immersion in liquid solvents or in detergent solutions is a suitable procedure. Hot alkali detergents are widely used for degreasing. All cleaning solutions must be thoroughly removed before soldering. Residues from hard-water rinses can interfere with soldering. Material compatibility is a prime consideration. Pickling. The purpose of pickling, or acid cleaning, is to remove rust, scale and oxides, or sulfides from the metal in order to provide a clean surface for soldering. All of the inorganic acids (that is, hydrochloric, sulfuric, phosphoric, nitric, and hydrofluoric acids), singly or mixed, will fulfill this function, although hydrochloric and sulfuric acids are the most widely used. The pieces should be washed thoroughly in hot water after pickling and then dried as quickly as possible. Mechanical cleaning includes the following methods:

• • • • •

GRIT AND SHOTBLASTING MECHANICAL SANDING AND GRINDING FILING AND HAND SANDING CLEANING WITH STEEL WOOL WIRE BRUSHING AND SCRAPING

For best results, cleaning should extend beyond the joint area. Shot or steel gritblasting is often effective and is preferable to sandblasting because it avoids the embedding of silica particles on the surface, which interferes with the flow of solder. Mechanical cleaning is not recommended for softer metals, such as copper. Precoating. The coating of the base metal surfaces with a more-solderable metal or alloy prior to the soldering operation can facilitate soldering. Coatings of tin, copper, silver, cadmium, iron, nickel, and the alloys of tin-lead, tin-zinc, tincopper, and tin-nickel are used for this purpose. The advantages of precoating include the following: • •

SOLDERING IS MORE RAPID AND UNIFORM. STRONG ACIDIC FLUXES CAN BE AVOIDED DURING SOLDERING.

The precoating of metals that have tenacious oxide films (for example, aluminum, aluminum bronzes, highly alloyed steels, and cast iron) is almost mandatory. The precoating of steel, brass, and copper, although not entirely essential, is of great value in some applications. A number of different methods are used to precoat the metal surfaces. Solder or tin can be applied using a soldering iron or an abrasive wheel, ultrasonic soldering, immersion in molten metal, electrodeposition, or chemical displacement. Hot dipping can be accomplished by fluxing and dipping the parts in molten tin or solder. Often, small parts are initially placed in wire baskets and then cleaned, fluxed, and dipped in the molten metal. Finally, they are centrifuged to remove the excess metal. Coating by hot dipping is applicable to carbon steel, alloy steel, cast iron, copper, and certain copper alloys. Prolonged immersion in molten tin or solder should be avoided to prevent the excessive formation of intermetallic compounds at the coating/base-metal interface. Electrodeposition. Precoating by electrodeposition can be accomplished in stationary tanks, in conveyorized plating units, or in barrels. This method is applicable to all steels, copper alloys, and nickel alloys. The coating metals are not limited to tin and solder. Copper, cadmium, silver, precious metals, nickel, iron, and alloy platings such as tin-copper, tinzinc, and tin-nickel are also commonly used.

Certain combinations of electrodeposited metals (that is, duplex coatings), where one metal is plated over another, are becoming more popular as an aid to soldering. A coating of 0.005 mm (0.0002 in.) of copper plus 0.0075 mm (0.0003 in.) of tin is particularly useful for brass. The solderability of aluminum is enhanced by a coating of 0.013 mm (0.0005 in.) of nickel, followed by 0.008 mm (0.0003 in.) of tin or by a combination of zincate (zinc), 0.005 mm (0.0002 in.) copper, and tin. An iron plating followed by tin plating is extremely useful over a cast iron surface. Immersion coatings (that is, chemical displacement coatings) of tin, silver, or nickel can be applied to some of the

common base metals. These coatings are usually very thin and generally have a poor shelf life. The shelf life of a coating is defined as its ability to withstand storage conditions without impairing solderability. Hot tinned and brightened electrotin coatings have an excellent shelf life. Inadequate thicknesses of electrotinned or immersion-tinned coatings have a limited shelf life. Tin or solder coating thicknesses ranging from 0.003 to 0.008 (0.0001 to 0.0003 in.) are recommended to ensure maximum solderability after prolonged storage.

Reference cited in this section

19. J. BROUS, EVALUATION OF POST-SOLDER FLUX REMOVAL, WELD. J., VOL 54 (NO. 12), 1975, P 444S-448S

Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Soldering in Electronic Applications This section addresses some general considerations in the soldering of electronic devices. More detailed information is provided in the article "Soldering in Electronic Applications," in this Volume. Substrates for Electronic Component Applications A substrate is the platform on which the electrical/electronic circuitry is built. Many different types of substrates are used in the electronics field. Early substrates were sheets of material to which insulated, stand-off terminals were attached by bolting or riveting. The circuitry was then obtained by point-to-point wiring between these terminals. Although this approach is still used, the most common substrate in use today has circuitry that is laid on it in planar form, with the components already attached. Attachment is normally accomplished through the soldering process. The most common form of this substrate is the printed wiring board (PWB). A substrate can actually have numerous forms: • • •

CIRCUITRY CAN BE LAID ON ONE SIDE OF THE SUBSTRATE OR ON BOTH SIDES. CIRCUITRY CAN BE NOT ONLY ON THE OUTER SURFACES OF THE SUBSTRATE, BUT IN LAYERS WITHIN THE SUBSTRATE MATERIAL (MULTILAYERING). THE SUBSTRATE CAN BE RIGID OR FLEXIBLE.

There are many methods of applying the conductive patterns that make up the circuitry and each has its own specific soldering solutions. Many different materials are used for substrates. Substrates used for electrical/electronic applications must be insulators to allow the conductive circuitry to function in the manner of a network made from separate insulated wires. Common substrate materials for this application include ceramics and organic laminates. Metals can also be used as substrates if they are coated with a substantial insulating material that can withstand the rigors of assembly. Ceramic Substrates. Many types of ceramics can be used as substrates, but the most common is a high-fired alumina

that is either 96 or 99% Al2O3. The 96% Al2O3 material is usually employed when thick-film materials are placed on the surface to form the circuitry. The 99% Al2O3 material is most often used for the evaporation or sputtering of thin materials. The materials that are laid down for terminal and component bonding pads are usually gold (or alloys high in gold), silver, and copper. Because these metals are laid down in thin layers, a soldering problem called leaching occurs. Leaching is the migration of the components of a substrate into a working solution. The process must be controlled because soldering results in bonding between the solder and the metals being attached. If the soldering process is not controlled, it can literally dissolve the metallization from the surface of the substrate. A gold-platinum alloy is more resistant to leaching than pure gold. o

Thin-film (10 to 500 nm, or 100 to 5000 A ) deposits are more susceptible to leaching than are thick-film (10 to 25 μm, or 400 to 1000 μin.) deposits, because the thin layer is more susceptible to being consumed by the leaching. Therefore, the time at temperature and the maximum temperature become very important. In leaching, temperature is more important than time. Significantly, more leaching will be generated by a small increase in temperature than would be produced with even a moderate increase in time. Composite Laminates. The most popular composites are the copper-clad laminates used for PWB substrates. The metal cladding is usually copper, but could be nickel or some other metal. Copper cladding can be either a plated material or wrought rolled foil. From a soldering standpoint, the rolled material is preferable, but it is more costly and more difficult to bond to the laminate material. Many PWBs are passed through numerous heat cycles during the manufacturing process. In addition, the use of multilayer designs is increasing dramatically. Plated foil, contaminated with codeposited organics,

can become brittle at elevated temperatures. Hence, foil coppers are available in several classes, depending on their application. Printed Wiring Laminate Materials. The laminate upon which the metal cladding is applied is available in many composite forms. Originally, layers of paper or cloth impregnated with a phenolic resin and designated XP, XXP, or XXXP were used. The latter became the standard in the mid-1950s.

The selection of organic-based laminate materials as electronic substrates has certain limitations, the most common of which is temperature. In addition, soldering must not degrade the material by burning or charring. Laminate selection should depend on the application of the electronic hardware. A glass-epoxy laminate material known as FR-4 is commonly used in consumer and industrial electronics. The "FR" designation indicates that the material is fire resistant. Metal Substrates. There are some applications in which metals can be used as base substrates. Because the presence of

an insulating material is required, the metal must be coated with a suitable material (for example, porcelain enamel). Solder Application Because of the small distances between components, when soldering electronics, it is often necessary to measure carefully the amount of solder that must be applied to form a satisfactory joint. It is not only a matter of providing a sufficient amount of solder for a good fillet, but it is also important to not have excess solder in the joint. This can sometimes be accomplished by plating the area to be soldered to a predetermined solder thickness, so that the solder forms a good fillet when it is reflowed. This is often the method used to accomplish "flip-chip" bonding to produce tape automated bonding (TAB) assemblies (Ref 20). Solder can also be apportioned in proper amounts to individual joints by using preforms. In the case of surface-mount technology (SMT), the solder is often screened onto the substrate in the form of solder paste or it is dispersed in small quantities in precise locations by using tiny nozzles or tubes. Preforms. When a square or round lid is attached to a device, but solder is not desired in the enclosed area, the solder

application is often accomplished by using squares or annular rings that are punched from solder sheet. The solder quantity is usually sufficient to fill the asperities in the surfaces. The thickness and the lateral dimensions are determined by the location and the amount of solder needed to make a satisfactory joint. The solder preform is placed in the space between the parts to be joined, and then this assembly is subjected to sufficient heat to fuse the solder and hermetically seal the assembly. The entire body is then cooled. If the parts were clean and the proper amount of solder was supplied, a good solder joint results. Heat for soldering can be applied in a number of different ways. In automated processing, the parts to be soldered are often passed through the hot zone of a furnace on a belt. This makes it possible to provide adequate preheating of the parts, the proper maximum temperature, and a gentle cooling cycle to minimize stresses. This method is fast and allows easy control of the measured amounts of solder to precise locations in a variety of geometric shapes. However, the cost of the preforms is a drawback. Solder Pastes. The rapid development of SMT processing, which can involve very dense populations of devices and

leads, necessitated a method to precisely place controlled amounts of solder on a substrate, either for lead attachment or for bonding of the device to attain more-effective heat transfer. Solder pastes that are applied by screening or printing through a mask or by application through a fine tube can fulfill these needs. The screening method, although ancient, has been modernized by the development of new high-precision screens, solder pastes, and equipment for screening and curing the screened materials. Solder paste also can be applied through a metal mask, or stencil. Although the location of the solder is not as precise as that obtained with screening, the metal masks are more durable and provide better control of solder thickness as it is applied. Another paste placement method is to use fine tubes or nozzles to deliver the paste to precise locations. This method is ideal for placing discrete dots of paste, rather than covering large areas.

Paste Composition. A solder paste must contain solder particles of the desired composition to yield the melting and

joining requirements of the application. The metal particles must be sized to provide good screenability and to avoid excessive oxide content. The shape of the particles is important. Spherical particles are lower in oxide content because there is less surface area and they are able to flow more freely in screening applications. A paste must also contain a vehicle to control its rheological, or flow, properties; its drying behavior; its surface activity (wetting behavior); and the solubility of the residues left after reflow. The vehicle usually consists of a rosin (pine derivative) or a synthetic resin, solvents to adjust drying time and temperature, modifiers to control rheological behavior, and an activator, or flux, to remove oxides from the solder and to protect the liquefied solder from oxidation. Paste Properties. The paste must set and hold its shape and location after screening so that subsequent processing

maintains the solder in the desired locations and amounts. The assembly is then subjected to heat and the solder is melted and joined to the part being soldered. This reflow can be conducted either as the final step in soldering or as a separate step. The parts must generally be cleaned after soldering. Solder resists enable control over the PWB areas that will be wet by the solder. The resist, or mask, is an organic

coating placed selectively on the board. Only the uncoated areas are exposed to and are wettable by the solder. The higher interconnect densities now being attained by SMT processing make solder resists a valuable method. Solder resists can be either permanent or temporary, depending on the application. The permanent resists must resist not only the solder, but any fluxes or solvents applied later in the manufacturing sequence. Compatibility with the process and the circuit materials is a consideration, as is the ability to withstand soldering temperatures. In addition, the solder resist offers the circuits some protection against dirt and contamination. Commonly used resist materials are polystyrene, epoxides, and acrylics, although more-costly materials, such as polyimides, are sometimes used because of their greater thermal resistance. Screened Materials. Solder masks are commonly applied by screening. The films are subsequently dried and cured

thermally. The elevated temperatures involved in the curing cycle can cause oxidation of the conductors on the PWB, resulting in a loss of solderability. Lower curing temperatures are desirable. Some of the resists are cured by exposure to ultraviolet light, rather than by heating. This approach is very fast, involves less emission of volatile pollutants, and does not cause oxidation of the metal surfaces. Photographic Films. A second method of applying solder resists is to use sheets of photosensitive material laminated to

the surface of the circuit. Although these are more expensive than screened materials, the photographic films enable much higher pattern resolution. Dry film resists are much stronger and more resistant to damage during soldering and cleaning operations. The use of these photosensitive materials increases the complexity of the process. Temporary solder masks are often used as production aids, offering local protection from solder at locations such as: • • •

CONTACT PADS THAT HAVE BEEN GOLD PLATED PLATED-THROUGH HOLES THAT MUST BE SOLDERED IN A LATER OPERATION PADS THAT MUST BE KEPT FREE OF SOLDER

Adhesive solder-mask tapes that specifically protect edge connectors on PWBs are also available, and numerous mechanical approaches to limiting solder contact exists (Ref 18).

References cited in this section

18. PROC. IEEE 41ST ELECT. COMP. & TECH. CONF., IEEE, 1991 20. J.H. LAU, HANDBOOK OF TAPE AUTOMATED BONDING, VAN NOSTRAND REINHOLD, 1992 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Soldering Process Parameters The parameters that affect wetting and spreading phenomena include: • • • • •

TEMPERATURE TIME VAPOR PRESSURE METALLURGICAL AND CHEMICAL NATURE OF THE SURFACES GEOMETRY OF THE SOLID

The manipulation of each parameter can result in some control of the wetting and soldering processes. Temperature is important to wetting and spreading for several reasons. Higher temperatures result in greater atomic

activity and provide some of the energy needed to overcome surface barriers. Higher temperatures also increase reaction rates exponentially. Temperature determines phase relationships for a given alloy composition, whether for the formation of a solid solution or for that of an intermetallic compound. The fluidity of the liquid solder is also increased with increasing temperature. In addition, higher temperatures promote the formation of oxides, which can be troublesome when soldering. Time. Wetting is time dependent because of the rate of kinetics of the solder-substrate interaction, the flux action, and the

heat conduction of the substrate. The soldering process is time sensitive, because time is needed to provide sufficient opportunity for the solder to wet, penetrate, or "wick up" into the various areas of the substrate that must be incorporated into the solder joint. It is clear that measures of the time to initiate wetting and of the time it takes for spreading to occur are needed. Vapor pressure is considered to be negligible for lead-tin solders, but it can be a significant factor when working with

certain solders. Caution must be exercised when selecting and working with alloys that have appreciable vapor pressures. The chemical and metallurgical nature of a surface affects wetting and spreading in several ways. Alloy content, as

well as the fluidity of the liquid, both affect wetting and, thus, the rate of spreading. Grain boundaries are wet differently than the bulk of the metal. The formation of solid solutions or intermetallic compounds also can have a considerable effect on wetting and spreading. The presence of oxide skins will prevent wetting and spreading, because the solder generally will not wet oxides. Likewise, films or particles of organic matter will interfere with wetting. Oils and silicones are frequent offenders and usually require different types of cleaning materials for their removal than the types used for oxides. Surface Geometry. Because contact angle is a critical measure of wetting and spreading behavior, surface geometry is

an important factor in the control of spreading. Soldering artisans have long used a wire brush not only to clean the surface to be soldered, but also to aid the spreading of the solder by providing sharp grooves that act as small capillaries to assist wetting. Isotropic spreading is actually what is needed. Another condition of surface interaction that is sometimes encountered in soldering is dewetting, where a metal surface is initially wet by solder, but as the solder cools, its cohesive forces exceed the forces involved in wetting, and the solder balls up on the metal surface. This usually results from surface contamination of the metal or from the entrapment of foreign particles in the metal surface, leading to weak wetting or, possibly, only local wetting. Base-Metal Selection. A sound soldered joint is achieved by selecting and using the proper materials and processes.

Base metals are usually selected to achieve the specific property requirements of a component. These properties can include strength, ductility, electrical conductivity, weight, corrosion resistance, and others. When soldering is required, the solderability of the base materials should also be a selection factor. Both flux selection and surface preparation will be affected by the solderability of the base materials to be joined.

The solderability of metals and alloys is not simply a matter of chemical nobility, as might be supposed when regarding the good solderability of the noble metals, which do not readily form oxide or tarnish films. Although cadmium and tin both form oxides readily, they are considered easy to solder. On the other hand, chromium, nickel, and aluminum also form oxide films readily, but are difficult to solder. The difference is in the extremely adherent, protective nature of the oxides formed on chromium, nickel, and aluminum, compared with the oxides that form on tin and cadmium (Table 2).

TABLE 2 RELATIVE SOLDERABILITY OF SELECTED METALS AND ALLOYS

EASY TO SOLDER • • • • • • •

PLATINUM GOLD COPPER SILVER CADMIUM PLATE TIN SOLDER PLATE

LESS EASY TO SOLDER • • • • • •

LEAD NICKEL PLATE BRASS BRONZE RHODIUM BERYLLIUM COPPER

DIFFICULT TO SOLDER • • • •

GLAVANIZED IRON TIN-NICKEL NICKEL-IRON MILD STEEL

VERY DIFFICULT TO SOLDER • • • •

CHROMIUM NICKEL-CHROMIUM NICKEL-COPPER STAINLESS STEEL

MOST DIFFICULT TO SOLDER • •

ALUMINUM ALUMINUM BRONZE

NOT SOLDERABLE • •

BERYLLIUM TITANIUM

It should be noted that chromium, nickel, and aluminum are all soldered regularly with good results, but that special attention must be given to the selection of fluxes, which must be very active. In many cases, the use of active fluxes is either restricted or not allowed. Therefore, these hard-to-solder metals and alloys always require special consideration in order to provide reproducible soldering. Numerous tests and methods to measure solderability are available. The tendencies of solder to either wet or spread on a given material are critical when evaluating candidate soldering systems. When hand soldering on single, large-dimension bodies prevailed, the skilled artisan could adjust parameters while watching the results develop. In most cases, visual determinations of wetting and spreading characteristics provided a sufficient basis for process control. Today, operating parameters must be controlled very carefully, because the soldering of components and/or electronic assemblies can involve many joints that are soldered at one time, or joints that are often hidden, or dimensions that are extremely small, or situations where joint-to-joint accommodations cannot be made. This necessitates the use of automatic soldering systems. Moreover, conditions must be uniform over the whole structure being soldered. In this environment, it is essential that the processes be based on measurements of wetting and spreading behavior, so that the controlling parameters can be accurately set. Additional information is available in the articles "Brazeability and Solderability of Engineered Materials," "General Soldering," and "Soldering in Electronic Applications" in this Volume. Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Soldering Equipment The proper application of heat is of paramount importance in any soldering operation. The solder should melt while the surface is heated to permit the molten solder to wet and flow over the surface. The best heating method is therefore another important consideration. Soldering Irons. The traditional soldering tool is the soldering iron, or bit, with a copper tip than can be heated electrically, by direct flame, or in an oven. Because soldering is a heat-transfer process, the maximum surface area of the heated tip should contact the base metal. The solder itself should not be melted upon the tip of the iron when a joint is being made. To lengthen the usable life of a copper tip, a coating of solder-wettable metal, such as iron with or without additional coatings, is applied to the surface of the copper. The rate of dissolution of the iron coating in molten solder is substantially less than the rate for copper. The iron coating also shows less wear, oxidation, and pitting than uncoated copper.

Soldering irons are available in a large variety of sizes and designs, ranging from a small pencil to special irons or bits that weigh 2 kg (5 lb) or more. The selection of the iron depends on the application and the quantity of heat needed at the joint. The heat recovery time of the iron should be fast enough to keep up with the job. Regardless of the heating method, the tip performs the following functions: • • • • •

STORING AND CONDUCTING HEAT FROM THE HEAT SOURCE TO THE PARTS BEING SOLDERED STORING MOLTEN SOLDER CONVEYING MOLTEN SOLDER WITHDRAWING SURPLUS MOLTEN SOLDER BRINGING THE WORKPIECE JOINT AREA TO THE SOLDERING TEMPERATURE

The angle at which the copper tip is applied to the work is important in terms of delivering maximum heat. The flat side of the tip should be applied to the work to obtain the maximum area of contact. Flux-cored solders should not be melted on the soldering tip, because this destroys the effectiveness of the flux. The cored solder should be touched to the

soldering tip to initiate good heat transfer, and then the solder should be melted on the work parts to complete the solder joint (Ref 11). Additional information is available in the article "Iron Soldering" in this Volume. Torch Soldering (TS). The selection of a gas torch is controlled by the size, mass, and configuration of the assembly to be soldered. When fast soldering is necessary, a flame is frequently used. The flame temperature is controlled by the fuel mixture used. Fuel gas burned with oxygen gives the highest flame temperature possible with that gas. The highest flame temperatures are attained with acetylene. Lower temperatures are obtained with propane, butane, natural gas, and manufactured gas.

Multiple flame tips, or burners, often have shapes that are suitable to the work. They can be designed to operate on oxygen and fuel gas, compressed air and fuel gas, or Bunsen-type torches. When adjusting tips or torches, care should be taken to avoid adjustments that result in a "sooty" flame, because the carbon deposited on the work will prevent the flow of solder (Ref 11, 15). Detailed information is available in the article "Torch Soldering" in this Volume. Dip soldering (DS) utilizes a molten bath of solder to supply both the heat and solder required to produce the joints

(Fig. 9a). This method is useful and economical, because an entire unit comprising any number of joints can be soldered in one operation after proper cleaning and fluxing. Fixtures are required to contain the unit and maintain proper joint clearances during solder solidification. The soldering pot should be large enough so that, at a given rate of production, the units being dipped will not appreciably lower the temperature of the solder bath.

FIG. 9 SELECTED SOLDERING TECHNIQUES USED FOR HIGH-VOLUME PRODUCTION APPLICATIONS. (A) DIPPING ON A STATIC BATH. (B) WAVE SOLDERING SYSTEM. (C) CASCADE SOLDERING SYSTEM

Additional information is available in the article "Dip Soldering" in this Volume. Wave Soldering (WS). Electronic components are commonly soldered to PWBs by wave soldering (Fig. 9b). The solder is pumped out of a narrow slot to produce a wave or series of waves. This approach, which lends itself to automatic soldering, involves a continuously pumped crest of solder that peaks at the PWB. Another method is "drag," or planar, soldering. This method uses a bath, into which the board is automatically dipped with controlled conveyorization (Ref 21, 22).

In cascade soldering (Fig. 9c), the solder flows down a trough because of gravity and is returned by pump to the upper reservoir. Both of these wave solder systems are excellent in that a virtually oxide-free solder surface is presented to the part, and the flux and vapors are dislodged by the flow of the solder. Integrated wave soldering systems for printed wiring assemblies provide units that can apply the flux, dry and preheat the board, solder components, and clean the completed assembly. Some systems have special features that apply the flux by passing the assemblies through a wave, by spraying, by rolling, or by dipping. Several systems use oil mixed with the solder to aid in eliminating icicles and bridging between conductor paths. Another system features dual waves that flow in the opposite direction to the travel of the board. Additional information is available in the article "Wave Soldering" in this Volume. Induction Soldering (IS). The only requirement for a material that is to be induction soldered is that it be an electrical

conductor. The rate of heating depends on the induced current flow. The distribution of heat obtained with induction heating is a function of the induced current frequency. The higher frequencies concentrate the heat at the surface of the workpiece. Several types of equipment are available for induction heating: the vacuum tube oscillator, the resonant spark gap, the motor-generator unit, and solid-state units. Induction heating is generally applicable for soldering operations, with the following requirements and attributes: • • • • •

LARGE-SCALE PRODUCTION APPLICATION OF HEAT TO A LOCALIZED AREA MINIMUM OXIDATION OF SURFACE ADJACENT TO THE JOINT GOOD APPEARANCE AND CONSISTENTLY HIGH JOINT QUALITY SIMPLE JOINT DESIGN THAT LENDS ITSELF TO MECHANIZATION

The induction technique requires that the parts being joined have clean surfaces and that joint clearances be maintained accurately. High-grade solders are generally required to obtain rapid spreading and good capillary flow. Preforms often afford the best means of supplying the correct amount of solder and flux to the joint. When induction soldering dissimilar metals (particularly joints composed of both magnetic and nonmagnetic components), attention must be given to the design of the induction coil in order to bring both parts to approximately the same temperature. Additional information is available in the article "Induction Soldering" in this Volume. Furnace Soldering. There are many applications, especially in high-production soldering, where a furnace will produce

consistent and satisfactory soldering. Although this method is not widely used, furnace heating should be considered under the following circumstances:

• • •

WHEN ENTIRE ASSEMBLIES CAN BE BROUGHT TO THE SOLDERING TEMPERATURE WITHOUT DAMAGE TO ANY OF THE COMPONENTS WHEN PRODUCTION IS SUFFICIENTLY LARGE TO ALLOW EXPENDITURES FOR JIGS AND FIXTURES TO HOLD THE PARTS DURING THE SOLDERING PROCESS WHEN THE ASSEMBLY IS COMPLICATED IN NATURE, MAKING OTHER HEATING METHODS IMPRACTICAL

The proper clamping fixtures are very important during furnace soldering. Movement of the joint during solder solidification can result in a poor joint. Another important consideration in furnace soldering is flux selection. Rosin and organic fluxes are subject to decomposition when maintained at elevated temperatures for an extended period of time. When either rosin or organic flux is used, the part must be brought rapidly to the liquidus temperature of the solder. It is sometimes beneficial to dip the parts in a hot flux solution before placing them in the oven. When using rosin-base flux, it is generally necessary to use a solder with a tin content of 50% or more. The reducing atmosphere used in the furnace does not allow joints to be made without flux, because the temperatures at which these atmospheres become reducing are far above the liquidus temperature of the solders. The use of inert atmospheres will prevent further oxidation of the parts, but flux must be used to remove the oxide that is already present. The furnaces should be equipped with adequate temperature controls because solder flow begins at approximately 45 to 50 °C (80 to 90 °F) above the solder liquidus temperature. The optimum condition exists when the heating capacity of the furnace is sufficient to heat the parts rapidly to the liquidus temperature of the solder. Detailed information is available in the article "Furnace and Infrared Soldering" in this Volume. Resistance soldering (RS) involves placing the workpiece either between a ground and a movable electrode or

between two movable electrodes as part of an electrical circuit. Heat is applied to the joint both by the electrical resistance of the metal being soldered and by conduction from the electrode, which is usually carbon. Production assemblies can utilize multiple electrodes, rolling electrodes, or special electrodes, depending on which setup offers the most advantageous soldering speed, localized heating, and power consumption. Resistance soldering electrode bits generally cannot be tinned, and the solder must be fed directly into the joint. Additional information is available in the article "Resistance Soldering" in this Volume. Infrared Soldering (IRS). Optical soldering systems that focus infrared light (radiant energy) on the joint by means of

a lens are available. Lamps with power ratings that range from 45 to 1500 W (140 to 4700 Btu/h) can be used for different application requirements. The devices can be programmed through a silicon-controlled rectifier (SCR) power supply with an internal timer. The most common sources of infrared heating for soldering applications are heated filaments. The quartz-iodine tungstenfilament lamp is widely used because it is very stable and reliable over a wide range of temperatures. In general, infrared soldering systems are simple and inexpensive to operate. One of the most critical operating parameters is surface condition. Variations in the condition of the solder surface can be compensated for, to some extent, by adjusting the heating power. Advantages are process repeatability, ability to concentrate or focus the energy with reflectors and lenses, economy of operation, and absence of contact with the workpiece. Detailed information is available in the article "Furnace and Infrared Soldering" in this Volume. Ultrasonic Soldering. This process uses a transducer as the source of ultrasonic energy. The transducer is energized in a bath of molten solder, and sound waves are coupled between the transducer and the workpiece, allowing the oxides in the base metal to be disrupted so that the solder melts the base metal. Sound waves are transmitted throughout the base metal, permitting wetting to occur on surfaces that are "blind" to the source. Ultrasonic soldering is also used to apply solderable coatings on difficult-to-solder metals.

Additional Soldering Processes. In addition to the above-mentioned processes, hot-gas soldering, spray-gun soldering,

and condensation soldering are also used. Hot gas soldering utilizes a fine jet of inert gas, heated to above the liquidus of the solder. The gas acts as a heat-

transfer medium and as a blanket to reduce access of air at the joint. Additional information is available in the article "Hot Gas Soldering" in this Volume. Spray gun soldering is a heating method used when the contour of the part to be soldered is difficult to follow with

either a wiping or drop method or when the part is placed in the assembly in such a way that the solder cannot be applied after the parts are assembled. Gas-fired or electrically heated guns are available. Each type is designed to spray molten or semimolten solder on the work from a continuously fed solid solder wire. Soldering guns use either propane mixed with oxygen or natural gas mixed with air to heat and to spray a continuously fed solid solder wire of approximately 3.2 mm (0.12 in.) diameter. About 90% of the solder wire is melted by the flame of the gun. The solder contacts the workpiece in a semiliquid form. The workpiece then supplies the balance of the heat required to melt and flow the solder. Adjustments can be made within the spray gun to control the solder spray. Vapor-phase soldering (also known as condensation soldering) utilizes the latent heat of vaporization of a condensing

saturated liquid to provide the heat required for soldering. A reservoir of saturated vapor over a boiling liquid provides a constant controlled temperature with rapid heat transfer. This method is useful for large assemblies, as well as for temperature-sensitive parts. Detailed information is available in the article "Vapor-Phase Soldering" in this Volume.

References cited in this section

11. R.W. WOODGATE, THE HANDBOOK OF MACHINE SOLDERING, 2ND ED., JOHN WILEY & SONS, 1988 15. E. LIEBERMAN, MODERN SOLDERING AND BRAZING TECHNIQUES, BUSINESS NEWS PUBLISHING CO., 1988 21. P.J. BUD, MASS PRODUCTION TECHNIQUES USING THE PRINCIPLE OF WAVE SOLDERING, WELD. J., VOL 52 (NO. 7), JULY 1973, P 431-439 22. P.J. BUD AND D.A. ELLIOT, WAVESOLDERING AT ELEVATED TEMPERATURES, WELD. J., VOL 53 (NO. 2), FEB 1974, P 79-87 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Quality Control Table 3 lists that are commonly used to evaluate the solderability properties of selected soldered components.

TABLE 3 TEST STANDARDS USED TO EVALUATE SOLDERABILITY

BASIC METHOD DIP AND LOOK

TEST STANDARD ANSI/IPC-S-804 (BOARDS) ANSI/IPC-S-805 (COMPONENTS)

IEC 68-2-20 MIL-STD-202, 750, 883 EIA RS-186-9E ROTARY DIP ANSI-IPC-S-804 (BOARDS) IEC 68-2-20 BS4025 (BRITISH STANDARD) WETTING BALANCE ANSI-IPC-S-805 (COMPONENTS) IEC-68-2-20 MIL-STD-883 GLOBULE TEST ANSI-IPC-S-805 (COMPONENTS) IEC 68-2-20 BS 2011 (BRITISH STANDARD) DIN 40046 (GERMAN STANDARD) MENISCUS RISE NO STANDARDS TO DATE TIMED SOLDER RISE ANSI/IPC-S-804 (BOARDS) Dip Test. The most useful test for assessing solderability is the dip test, because of its accuracy and simplicity. It is

conducted with the additions of cams and timers so that standardized conditions are established. The dip test is a reasonable simulation of practical soldering conditions and starts with a test specimen that has been dipped in a mild flux, as is usually the case with soldering tests. The dip test involves immersing the test specimen into a clean solder bath, waiting an appropriate length of time, and then withdrawing the specimen from the bath. In most cases, the test is automated to ensure reproducibility. The results are evaluated by means of visual inspection and comparison with a set of standards. Thus, the dip test is operator sensitive, somewhat subjective, and of limited value, because it is not quantitative. Rotary Dip Test. This test was devised to provide a better simulation of the dynamic conditions that exist in machine

soldering operations. It is used to evaluate the solderability of PWBs, and involves bringing the specimen into moving contact with the molten solder for a specified period of time. This is a "go/no-go" test that requires complete wetting of the specimen surface within a set time, and is thus dependent on visual inspection of the test specimen. Wetting Balance (Meniscograph). Another common test employs the wetting balance, which gives a numerical result.

The equipment consists of a balance that records the forces exerted on a fluxed test specimen during a controlled dipping cycle in a clean solder bath. From the values obtained, it is possible to compare the rate of wetting, as well as the forces involved for various flux formulations. The speed of wetting is an especially important factor in the wave soldering process, where automation requires a timebased response of flux and solder as the work moves through the process machine at a constant speed. The wetting balance test cannot be used on finished, etched printed circuits or on single-sided circuits. It is often used on specimens cut from double-sided clad materials prior to etching. Although this test gives numerical results, they are not strictly quantitative. However, the results do provide the basis for a meaningful comparison among fluxes. Additional Tests. There are several other tests that provide practical information for special purposes, such as the

spreading test and the globule test. These tests are sometimes used to evaluate incoming materials or to aid in the selection of materials. They are not universally employed. The spreading test is used to provide a numerical rating for both fluxes and solders. It is described in detail in MIL-F-

14256, "Flux, Soldering, Liquid (Rosin Base)," and in QQ-S-571, "Solder, Electronic," which are national-level specifications. The globule test was once widely applied, but is now limited to being used to determine the wetting of component leads

and wires (Ref 24). Additional information is available in the article "Evaluation and Quality Control of Soldered Joints" in this Volume.

Reference cited in this section

24. IEC-68-2-20, INTERNATIONAL ELECTROCHEMICAL COMMISSION, GENEVA Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

Future Outlook The quality soldered joints obtained by automation have resulted in the widespread use of automated equipment throughout the electronics industry, as well as throughout the world (Ref 25). When compared to processes such as adhesive bonding, welding, brazing, or mechanical joining, soldering offers many advantages. Automated soldering equipment for wave or planar soldering can be easily and economically installed. The high reliability that can be achieved with soldering results in a quality end product. Soldering offers the additional benefit of providing sequential assembly, a common practice in the electronics industry.

Reference cited in this section

25. REFLOW SOLDERING, WELD. J., VOL 52 (NO. 1), JAN 1973, P 22-30 Fundamentals of Soldering Mel M. Schwartz, Sikorsky Aircraft

References

1. J. WOLTERS, ZUR GESEHICHTE DER LOTTECHNIK, DEGUSSA, WEST GERMANY, 1977 2. H.H. MANKO, SOLDERS AND SOLDERING, 2ND ED., MCGRAW-HILL, 1964 3. R.J. KLEIN WASSINK, SOLDERING IN ELECTRONICS, ELECTROCHEMICAL PUBLICATIONS LTD., 1989 4. L.P. LAMBERT, SOLDERING FOR ELECTRONIC ASSEMBLIES, MARCEL DEKKER, 1987 5. SOLDERING MANUAL, AWS, 1978 6. A.D. ROMIG, JR., Y.A. CHANG, J.J. STEPHENS, ET AL., PHYSICAL METALLURGY OF SOLDERSUBSTRATE REACTIONS, SOLDER MECHANICS, THE MINERALS, METALS, & MATERIALS SOCIETY, 1991, P 30-32 7. "SOFT SOLDER ALLOYS--CHEMICAL COMPOSITIONS AND FORMS," ISO 9453, 1ST ED., 1990 8. "SOFT SOLDERING FLUXES--CLASSIFICATION AND REQUIREMENTS--PART I: CLASSIFICATION, LABELING AND PACKAGING," ISO 9454-1, 1990 9. J.F. SHIPLEY, INFLUENCE OF FLUX, SUBSTRATE AND SOLDER COMPOSITION ON SOLDER WETTING, WELD. J., VOL 54 (NO. 10), OCT 1975, P 357S-362S 10. H.H. MANKO, SOLDERING FLUXES--PAST AND PRESENT, WELD. J., VOL 52 (NO. 3), MARCH 1973, P 163-166 11. R.W. WOODGATE, THE HANDBOOK OF MACHINE SOLDERING, 2ND ED., JOHN WILEY & SONS, 1988

12. T. THOMPSON, CONDENSATION/VAPOR-PHASE REFLOW SOLDERING, ASS. ENG., JUNE 1977, P 44-47 13. T.Y. CHU, A GENERAL REVIEW OF MASS SOLDERING METHODS, INS. CIRC., NOV 1976 14. R.C. PFAHL, JR., J.C. MOLLENDORF, AND T.Y. CHU, CONDENSATION SOLDERING, WELD. J., VOL 54 (NO. 1), 1975 15. E. LIEBERMAN, MODERN SOLDERING AND BRAZING TECHNIQUES, BUSINESS NEWS PUBLISHING CO., 1988 16. "SOLDER METAL SPECIFICATION," ASTM B 32, ANNUAL BOOK OF ASTM STANDARDS, PART B, ASTM 17. D.R. FREAR, W.B. JONES, AND K.R. KINSMAN, ED., SOLDER MECHANICS, TMS, 1991, P 30-41 18. PROC. IEEE 41ST ELECT. COMP. & TECH. CONF., IEEE, 1991 19. J. BROUS, EVALUATION OF POST-SOLDER FLUX REMOVAL, WELD. J., VOL 54 (NO. 12), 1975, P 444S-448S 20. J.H. LAU, HANDBOOK OF TAPE AUTOMATED BONDING, VAN NOSTRAND REINHOLD, 1992 21. P.J. BUD, MASS PRODUCTION TECHNIQUES USING THE PRINCIPLE OF WAVE SOLDERING, WELD. J., VOL 52 (NO. 7), JULY 1973, P 431-439 22. P.J. BUD AND D.A. ELLIOT, WAVESOLDERING AT ELEVATED TEMPERATURES, WELD. J., VOL 53 (NO. 2), FEB 1974, P 79-87 23. ELECTRONIC MATERIALS HANDBOOK, VOL 1, PACKAGING, ASM INTERNATIONAL, 1989 24. IEC-68-2-20, INTERNATIONAL ELECTROCHEMICAL COMMISSION, GENEVA 25. REFLOW SOLDERING, WELD. J., VOL 52 (NO. 1), JAN 1973, P 22-30 Introduction to Solid-State Welding Ray Dixon, Los Alamos National Laboratory

Solid-State Welding (SSW) processes are those that produce coalescence of the faying surfaces at temperatures below the melting point of the base metal being joined without the addition of brazing or solder filler metal. Pressure may or may not be applied. These processes involve either the use of deformation or of diffusion and limited deformation in order to produce high-quality joints between both similar and dissimilar materials. Dissimilar metal joints are necessary in applications that require a variety of material properties within the same component. For example, heat exchangers often require different types of stainless steels at each end, because of temperature-induced corrosion. Under laboratory conditions, dissimilar materials can be chosen based on physical or material properties that influence the phenomenon being studied. For whatever reason, and appropriate method of producing dissimilar metal joints can usually be determined (assuming it is even possible) by examining the phase diagram. If the diagram indicates difficulty in joining the materials (intermetallics, and so on), then a solid-state (nonmelting) process may be applicable. When a nonmelting process is chosen, it is only successful if a relatively strong joint is produced. There is considerable interest is quantifying joint strength, and extensive literature (Ref 1, 2, 3, 4, 5, 6, 7, 8, 9, 10, 11, 12, 13, 14) exists to help design and evaluate various specimen geometries. This experience includes not only techniques for conducting the tests, but analysis and interpretation of the data with respect to the physical properties of individual materials. The study of dissimilar welds necessarily leads to studies of the interfacial region, because it is the most likely site of part failures. In a two-component system (materials A and B), interfacial failure can be a result of the individual properties of either A and B, but they are by necessity synergistic. These synergisms and the resulting complex interactions that they produce are what is significant. From an engineering and processing standpoint, an interfacial "cookbook" that provides

all the answers would be ideal. However, this capability is not yet available, and, as explained in this Section, researchers are still far from reaching that goal. This is the case even for simple systems, such as metal pair that form solid solutions. The opportunities for understanding interfacial behavior come from studies conducted at the atomic level (for example, using scanning tunneling microscopy, transmission electron microscopy, first principles energy band calculations, and embedded atom potential calculations) with appropriate correlations to both micromechanics and a macroscopic level, and vice versa. The computational and experimental tools that begin to provide insight at these levels have only recently been developed. Because the interest in interfaces is so diverse (all combinations of metals, ceramics, semiconductors, glasses, and conducting polymers), the advances in interfacial science encompass the creative efforts of a large number of people and organizations. Thus, the quantity of available literature seems overwhelming, and perhaps, disjointed. It will become apparent that there is no preferred approach to the treatment of dissimilar welds or welding. It really depends on the interests and objectives of each engineer. For example, if you know that what you are doing actually works (that is, produces a joint) and are searching for understanding, then this Section of the Handbook will guide you to the appropriate technical literature. If you have an idea and need guidance for bonding, testing, processing, or interpreting the data, then this Section should also provide that direction. Lastly, if the reader wants state-of-the-art research and development topics, then those can be found in this Section, as well. This should provide the reader with a comprehensive overview of the welding of dissimilar metals. Therefore, this Section is concerned with the fundamentals of welding and joining materials via the application of a nonmelting process. The specific processes usually associated with this technology are also discussed. Before considering nonmelting processes in detail, however, a brief discussion of adhesion is presented. This is followed by articles that discuss the fundamentals of the major solid-state welding techniques.

References

1. P.G. CHARALAMBIDES, J. LUND, A.G. EVANS, AND R.M. MCMEEKING, "A TEST FOR DETERMINING THE FRACTURE RESISTANCE OF BIMATERIAL INTERFACES," REPORT M87-4, MATERIALS DEPARTMENT, COLLEGE OF ENGINEERING, UC SANTA BARBARA 2. Z. SUO AND J.W. HUTCHINSON, "INTERFACE CRACK BETWEEN TWO ELASTIC LAYERS," REPORT MECH-118, DIVISION OF APPLIED SCIENCES, HARVARD UNIVERSITY, MARCH 1988 3. N.P. O'DOWD, C.F. SHIH, AND M.G. STOUT, TEST GEOMETRIES FOR MEASURING INTERFACIAL FRACTURE TOUGHNESS, FOR INT. J. SOLIDS STRUCT., VOL 29, 1992, P 571-589 4. J.W. HUTCHINSON AND Z. SUO, MIXED MODE CRACKING IN LAYERED MATERIALS, ADV. IN APPL. MECH., VOL 28, J.W. HUTCHINSON AND T.Y. WU, ED., ACADEMIC PRESS, 1990 5. H.M. JENSEN, MIXED MODE INTERFACE FRACTURE CRITERIA, ACTA METALL. MATER., VOL 38 (NO. 12), 1990, P 2637-2644 6. R.L. BRADY, R.S. PORTER, AND J.A. DONOVAN, A BUCKLED PLATE TEST TO MEASURE INTERFACIAL TOUGHNESS IN COMPOSITES, J. MATER. SCI., VOL 24, 1989, P 4138-4143 7. M.-Y. HE AND J.W. HUTCHINSON, "KINKING OF A CRACK OUT OF AN INTERFACE," REPORT MECH-113, DIVISION OF APPLIED SCIENCE, HARVARD UNIVERSITY, FEB 1988 8. J.J. MECHOLSKY AND L.M. BAKER, A CHEVRON-NOTCHED SPECIMEN FOR FRACTURE TOUGHNESS MEASUREMENTS OF CERAMIC-METAL INTERFACES, CHEVRON-NOTCHED SPECIMENS: TESTING AND STRESS ANALYSIS, STP 855, J.H. UNDERWOOD, S.W. FREIMAN, AND F.I. BARATTA, ED., ASTM, 1984, P 324-336 9. W.W. GERBERICH, D.L. DAVIDSON, AND M. KACZOROWSKI, EXPERIMENTAL AND THEORETICAL STRAIN DISTRIBUTIONS FOR STATIONARY AND GROWING CRACKS, J. MECH. PHYS. SOLIDS, VOL 38 (NO. 1), 1990, P 87-113 10. J. QU AND Q. LI, INTERFACIAL DISLOCATION AND ITS APPLICATIONS TO INTERFACE CRACKS IN ANISOTROPIC BIMATERIALS, J. ELAST., VOL 26, 1991, P 169-195

11. A.N. STROH, DISLOCATIONS AND CRACKS IN ANISOTROPIC ELASTICITY, PHILOS. MAG., VOL 7, 1958, P 625-646 12. Q.Q. LI AND T.C.T. TING, LINE INCLUSION IN ANISOTROPIC ELASTIC SOLIDS, J. APPL. MECH., VOL 56, 1989, P 556-563 13. T.C.T. TING, EXPLICIT SOLUTION AND INVARIANCE OF THE SINGULARITIES AT AN INTERFACE CRACK IN ANISOTROPIC COMPOSITES, INT. J. SOLIDS STRUCT., VOL 22, 1986, P 965-983 14. M.D. PASHLEY AND D. TABOR, ADHESION AND DEFORMATION PROPERTIES OF CLEAN AND CHARACTERIZED METAL MICRO-CONTACTS, VACUUM, VOL 10-12, 1981, P 619-623 Introduction to Solid-State Welding Ray Dixon, Los Alamos National Laboratory

References

1. P.G. CHARALAMBIDES, J. LUND, A.G. EVANS, AND R.M. MCMEEKING, "A TEST FOR DETERMINING THE FRACTURE RESISTANCE OF BIMATERIAL INTERFACES," REPORT M87-4, MATERIALS DEPARTMENT, COLLEGE OF ENGINEERING, UC SANTA BARBARA 2. Z. SUO AND J.W. HUTCHINSON, "INTERFACE CRACK BETWEEN TWO ELASTIC LAYERS," REPORT MECH-118, DIVISION OF APPLIED SCIENCES, HARVARD UNIVERSITY, MARCH 1988 3. N.P. O'DOWD, C.F. SHIH, AND M.G. STOUT, TEST GEOMETRIES FOR MEASURING INTERFACIAL FRACTURE TOUGHNESS, FOR INT. J. SOLIDS STRUCT., VOL 29, 1992, P 571-589 4. J.W. HUTCHINSON AND Z. SUO, MIXED MODE CRACKING IN LAYERED MATERIALS, ADV. IN APPL. MECH., VOL 28, J.W. HUTCHINSON AND T.Y. WU, ED., ACADEMIC PRESS, 1990 5. H.M. JENSEN, MIXED MODE INTERFACE FRACTURE CRITERIA, ACTA METALL. MATER., VOL 38 (NO. 12), 1990, P 2637-2644 6. R.L. BRADY, R.S. PORTER, AND J.A. DONOVAN, A BUCKLED PLATE TEST TO MEASURE INTERFACIAL TOUGHNESS IN COMPOSITES, J. MATER. SCI., VOL 24, 1989, P 4138-4143 7. M.-Y. HE AND J.W. HUTCHINSON, "KINKING OF A CRACK OUT OF AN INTERFACE," REPORT MECH-113, DIVISION OF APPLIED SCIENCE, HARVARD UNIVERSITY, FEB 1988 8. J.J. MECHOLSKY AND L.M. BAKER, A CHEVRON-NOTCHED SPECIMEN FOR FRACTURE TOUGHNESS MEASUREMENTS OF CERAMIC-METAL INTERFACES, CHEVRON-NOTCHED SPECIMENS: TESTING AND STRESS ANALYSIS, STP 855, J.H. UNDERWOOD, S.W. FREIMAN, AND F.I. BARATTA, ED., ASTM, 1984, P 324-336 9. W.W. GERBERICH, D.L. DAVIDSON, AND M. KACZOROWSKI, EXPERIMENTAL AND THEORETICAL STRAIN DISTRIBUTIONS FOR STATIONARY AND GROWING CRACKS, J. MECH. PHYS. SOLIDS, VOL 38 (NO. 1), 1990, P 87-113 10. J. QU AND Q. LI, INTERFACIAL DISLOCATION AND ITS APPLICATIONS TO INTERFACE CRACKS IN ANISOTROPIC BIMATERIALS, J. ELAST., VOL 26, 1991, P 169-195 11. A.N. STROH, DISLOCATIONS AND CRACKS IN ANISOTROPIC ELASTICITY, PHILOS. MAG., VOL 7, 1958, P 625-646 12. Q.Q. LI AND T.C.T. TING, LINE INCLUSION IN ANISOTROPIC ELASTIC SOLIDS, J. APPL. MECH., VOL 56, 1989, P 556-563 13. T.C.T. TING, EXPLICIT SOLUTION AND INVARIANCE OF THE SINGULARITIES AT AN INTERFACE CRACK IN ANISOTROPIC COMPOSITES, INT. J. SOLIDS STRUCT., VOL 22, 1986, P 965-983 14. M.D. PASHLEY AND D. TABOR, ADHESION AND DEFORMATION PROPERTIES OF CLEAN AND CHARACTERIZED METAL MICRO-CONTACTS, VACUUM, VOL 10-12, 1981, P 619-623

Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Introduction IN ANY SOLID-STATE (NONMELTING) WELDING PROCESS, there are two primary areas of concern: Will the material bond and how strong is the bond? Literature that deals with the science and technology of material compatibility and bond strength is readily available. This article will review recent work in quantifying adhesion, bonding, and interfacial characterization and strength. Subsequent articles in this Section of the Handbook focus on the engineering principles associated with the formation of welds using these techniques. This discussion concentrates on metal-metal configurations, with only passing comments on metal-ceramic and metal/ceramic-semiconductor interfaces. This restriction is necessary to limit the length of this article. Wherever appropriate, applications of thin films on substrates will also be used. Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Adhesion Energy Surface energy, Γsurf, is defined as the increase in energy, ∆E, of a system divided by the area, A, caused by the creation of the surface:

ΓSURF = Σ∆EI/A

(EQ 1)

The atoms that contribute to the extra energy, ∆Ei, are usually within 10 lattice parameters or less of the surface (or interface) (Ref 1, 2, 3). A freshly created surface has a high surface energy, whereas a relaxed or reconstructed one is lower, by about 3 to 5%. At 0 K (-459 °F), surface energies vary about 15 to 25% for planes of various indices. The variation is less at higher temperatures, because of entropy. For example, the (111) surface of a face-centered cubic (fcc) structure has the lowest energy, whereas the (210) surface is about 15% higher (Ref 1, 2, 3). The ideal fracture energy, coh, which is used in estimating the Griffith criterion, is twice the surface energy:

ΓCOH (HKL) = 2 · ΓSURF (HKL)

(EQ 2)

where (hkl) are the indices of the plane. The actual energy consumed in fracture or crack propagation is composed of two terms: one is the Γcoh and the other is the plastic work consumed per unit area Γplastic (Ref 4, 5, 6, 7, 8). For ionic crystals, Γplastic is usually a few times Γcoh; for metals, Γcoh is about 1 to 4 J/m2 (2 × 10-5 to 1 × 10-4 cal/cm2); and Γplastic is about 1000 J/m2 (2 × 10-2 cal/cm2). For ceramic systems, Γcoh is about the same, but Γplastic is about 10 to 100 J/m2 (2 × 10-4 to 2 × 10-3 cal/cm2) (Ref 7). Γplastic and Γcoh are related, because when Γcoh is zero, no plastic work (Γplastic) will be consumed. The exact form of this dependence is not known, although some attempts have been made (Ref 4, 5, 6) to correlate the two terms.

References cited in this section

1. S.P. CHEN, D.J. SROLOVITZ, AND A.F. VOTER, COMPUTER SIMULATION ON SURFACES AND [001] TILT GRAIN BOUNDARIES IN NI, AL, AND NI3AL, J. MATER. RES., VOL 4, 1989, P 62-77 2. S.P. CHEN AND A.F. VOTER, RECONSTRUCTION OF THE (310), (210), AND (110) SURFACES IN FCC METALS, SURF. SCI. LETT., VOL 244, 1991, P L107-L112 3. S.P. CHEN, THEORETICAL STUDIES OF METALLIC INTERFACES, MATER. SCI. ENG. B, VOL 6, 1990, P 113-121

4. M.L. JOKL, V. VITEK, AND C.J. MCMAHON, JR., A MICROSCOPIC THEORY OF BRITTLE FRACTURE IN DEFORMABLE SOLIDS: A RELATION BETWEEN IDEAL WORK TO FRACTURE AND PLASTIC WORK, ACTA METALL., VOL 28, 1980, P 1479-1488 5. M.L. JOKL, V. VITEK, AND C.J. MCMAHON, JR., ON THE MICROMECHANICS OF BRITTLE FRACTURE: EXISTING VS INJECTED CRACKS, ACTA METALL., VOL 37, 1989, P 87-97 6. J. HACK, S.P. CHEN, AND D. SROLOVITZ, A KINETIC CRITERION FOR QUASI-BRITTLE FRACTURE, ACTA METALL., VOL 37, 1989, P 1957-1970 7. A. KELLY AND N.H. MACMILLAN, STRONG SOLIDS, 3RD ED., OXFORD SCIENCE PUBLICATIONS, 1986 8. V. VITEK, MICROMECHANISMS OF INTERGRANULAR BRITTLE FRACTURE IN INTERMETALLIC COMPOUND, J. PHYS. III, VOL 1, 1991, P 1085-1097 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Grain Boundary Energy The grain boundary energy, Γgb, is defined the same way as surface energy is, but it depends on the relative orientation of the adjacent grains (Ref 1):

ΓGB = Σ∆EI/AI (GB)

(EQ 3)

At misorientation angles of greater than 25°, Γgb usually reaches a plateau that characterizes random grain boundaries. Special grain boundaries, like Σ5 (210) or (310), tend to have slightly lower energies (Ref 1, 2, 3, 9). Twin boundary energies are even lower than special grain boundaries. For metals, Γgb usually ranges from 40 to 60% of the surface energy (Ref 10), whereas the twin energy is only about 5 to 15%. Grain boundary cohesive energy is defined as (Ref 1):

ΓCOH (GB) = ΓSURF (GRAIN A) + ΓSURF (GRAIN B) - ΓGB

(EQ 4)

For general and random grain boundaries, grain boundary cohesive energy is smaller. General grain boundaries tend to be weaker than the small angle boundaries.

References cited in this section

1. S.P. CHEN, D.J. SROLOVITZ, AND A.F. VOTER, COMPUTER SIMULATION ON SURFACES AND [001] TILT GRAIN BOUNDARIES IN NI, AL, AND NI3AL, J. MATER. RES., VOL 4, 1989, P 62-77 2. S.P. CHEN AND A.F. VOTER, RECONSTRUCTION OF THE (310), (210), AND (110) SURFACES IN FCC METALS, SURF. SCI. LETT., VOL 244, 1991, P L107-L112 3. S.P. CHEN, THEORETICAL STUDIES OF METALLIC INTERFACES, MATER. SCI. ENG. B, VOL 6, 1990, P 113-121 9. S.P. CHEN, STUDIES OF IRIDIUM SURFACES AND GRAIN BOUNDARIES, PHIL. MAG. A, VOL 65, 1992, TO BE PUBLISHED 10. L.E. MURR, INTERFACIAL PHENOMENA IN METALS AND ALLOYS, ADDISON-WESLEY, 1975 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Interfacial Energy By definition, the interfacial energy (Ref 2, 3) between two bulk materials (A and B) is the same as the grain boundary energy, except that there are two reference states:

ΓINTERFACE = [Σ∆EI (A) + Σ∆EJ(B)]/A

(EQ 5)

where ∆Ei(A) · [(∆Ej(B)] is the extra energy of atom i of A (B) type, compared to its bulk state of A (B) materials. The interfacial energy between similar metals tends to be small. For dissimilar materials, however, interfacial energies tend toward large positive values if they do not form compounds and large negative values if they do form compounds. The ideal adhesion energy is similar to grain boundary cohesion and is the difference between the two surfaces and the interfacial energies:

ΓADHESION = ΓSURFACE (A) + ΓSURFACE (B) - ΓINTERFACE (A/B)

(EQ 6)

References cited in this section

2. S.P. CHEN AND A.F. VOTER, RECONSTRUCTION OF THE (310), (210), AND (110) SURFACES IN FCC METALS, SURF. SCI. LETT., VOL 244, 1991, P L107-L112 3. S.P. CHEN, THEORETICAL STUDIES OF METALLIC INTERFACES, MATER. SCI. ENG. B, VOL 6, 1990, P 113-121

Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Theory of Adhesion Rose et al. (Ref 11) have calculated the bimaterial interfacial cohesion for many combinations of elements using band structure methods. They found a universal scaling relationship for the adhesion of different materials that extends beyond metal-metal interfaces to nonmetal systems. A few calculations are necessary to identify the adhesion behavior of the system and set the appropriate energy, length, and elastic constant scales. Plastic contributions are not treated in this formalism. The plastic work is more likely to be solved by incorporating the atomistic simulation and continuum analysis of Jokl et al. (Ref 4, 5) and Hack et al. (Ref 6). The tight binding method give qualitative trends as the energy band is filled for calculations using the elements in the periodic table (Ref 12, 13, 14). The correct trend seems to be captured by the band-filling argument pioneered in the original work of Friedel (Ref 15). For noble metals, a rigid band picture is preferred (Ref 14). Another useful tool is the Pettifor structure map. By using the Mendelev number for each of the 103 elements (Ref 13, 16), Pettifor found that alloying tendencies can be grouped into separate regions. If the interfacial chemistry and average Mendelev number are known, then the interfacial structure can be found from the Pettifor map. Extensive atomistic simulation work, using effective medium theory (Ref 17, 18), embedded atom potentials (Ref 19), or local volume potentials (Ref 1, 2, 3, 8, 9, 20, 21, 22, 23) had been very useful in describing the properties of many metals and compounds and their interactions with impurities such as boron and sulfur in Ni3Al and NiAl (Ref 20, 21). These atomic interaction descriptions include pair potential and many body interactions and have been obtained for systems that include Ni3Al, Cu3Au, NiAl, nickel, aluminum, iridium, gold, platinum, silver, copper, molybdenum, tantalum, iron, tungsten, niobium, beryllium, titanium, hafnium, zirconium, and rhenium (Ref 1, 2, 3, 9, 20, 21, 22, 23, 24, 25, 26, 27). These studies indicate that stoichiometry (Ref 1, 20, 21) and impurities (Ref 20, 21) can be crucial in determining ideal and realistic adhesion, including effects induced by boron and sulfur. They represent the fundamentals for future work,

which is expected to have significant impact on the theory using atomistic simulation under well-controlled situations for all conditions of stress, temperature, and chemical environments. These calculations should enable direct linking of theoretical studies and materials by design concepts, as applied in NiAl microalloyed with boron or iron (Ref 24, 25, 26, 27). DeBoer et al. (Ref 28) have developed excellent tools for screening bonding materials, based on two properties of the elements: electronegativity and electron density. One can use the Miedema formula for the adhesion of a bimetallic interface, based on the work function (closely related to the Pauling electronegativity), Φ , and the electron density, n, at the Wigner-Seitz radius, as described in Ref 28.

References cited in this section

1. S.P. CHEN, D.J. SROLOVITZ, AND A.F. VOTER, COMPUTER SIMULATION ON SURFACES AND [001] TILT GRAIN BOUNDARIES IN NI, AL, AND NI3AL, J. MATER. RES., VOL 4, 1989, P 62-77 2. S.P. CHEN AND A.F. VOTER, RECONSTRUCTION OF THE (310), (210), AND (110) SURFACES IN FCC METALS, SURF. SCI. LETT., VOL 244, 1991, P L107-L112 3. S.P. CHEN, THEORETICAL STUDIES OF METALLIC INTERFACES, MATER. SCI. ENG. B, VOL 6, 1990, P 113-121 4. M.L. JOKL, V. VITEK, AND C.J. MCMAHON, JR., A MICROSCOPIC THEORY OF BRITTLE FRACTURE IN DEFORMABLE SOLIDS: A RELATION BETWEEN IDEAL WORK TO FRACTURE AND PLASTIC WORK, ACTA METALL., VOL 28, 1980, P 1479-1488 5. M.L. JOKL, V. VITEK, AND C.J. MCMAHON, JR., ON THE MICROMECHANICS OF BRITTLE FRACTURE: EXISTING VS INJECTED CRACKS, ACTA METALL., VOL 37, 1989, P 87-97 6. J. HACK, S.P. CHEN, AND D. SROLOVITZ, A KINETIC CRITERION FOR QUASI-BRITTLE FRACTURE, ACTA METALL., VOL 37, 1989, P 1957-1970 8. V. VITEK, MICROMECHANISMS OF INTERGRANULAR BRITTLE FRACTURE IN INTERMETALLIC COMPOUND, J. PHYS. III, VOL 1, 1991, P 1085-1097 9. S.P. CHEN, STUDIES OF IRIDIUM SURFACES AND GRAIN BOUNDARIES, PHIL. MAG. A, VOL 65, 1992, TO BE PUBLISHED 11. J.H. ROSE, J.R. SMITH, F. GUINEA, AND J. FERRANTE, UNIVERSAL FEATURES OF THE EQUATION OF STATE OF METALS, PHYS. REV. B, VOL 29, 1984, P 2963-2969 12. T.B. MASSALSKI, ED., BINARY ALLOYS PHASE DIAGRAMS, ASM, 1986, P 2176-2181 13. D.G. PETTIFOR, ELECTRON THEORY OF METALS, PHYSICAL METALLURGY, R.W. CAHN AND P. HAASEN, ED., NORTH-HOLLAND, AMSTERDAM, 1983, P 73-152 14. J.S. FAULKNER, THE MODERN THEORY OF ALLOYS, PROG. MATER. SCI., VOL 27, 1982, P 1-187 15. J. FRIEDEL, TRANSITION METALS, ELECTRONIC STRUCTURE OF THE D-BAND, ITS ROLE IN THE CRYSTALLINE AND MAGNETIC STRUCTURES, THE PHYSICS OF METALS, J.M. ZIMAN, ED., CAMBRIDGE UNIVERSITY PRESS, 1969, P 340-408 16. D.G. PETTIFOR, NEW MANY-BODY POTENTIAL FOR THE BOND ORDER, PHYS. REV. LETT., VOL 63, 1989, P 2480-2483 17. J.K. NORSKOV AND N.D. LANG, EFFECTIVE-MEDIUM THEORY OF CHEMICAL BINDING: APPLICATION TO CHEMISORPTION, PHYS. REV. B, VOL 21, 1980, P 2131-2136 18. M.J. STOTT AND E. ZAREMBA, QUASIATOMS: AN APPROACH TO ATOMS IN NONUNIFORM ELECTRONIC SYSTEMS, PHYS. REV. B, VOL 22, 1980, P 1564-1583 19. M.S. DAW AND M.I. BASKES, EMBEDDED-ATOM METHOD: DERIVATION AND APPLICATION TO IMPURITIES, SURFACES, AND OTHER DEFECTS IN METALS, PHYS. REV. B, VOL 29, 1984, P 6443-6453 20. S.P. CHEN, A.F. VOTER, R.C. ALBERS, A.M. BORING, AND P.J. HAY, INVESTIGATION OF BORON EFFECT ON NI AND NI3AL GRAIN BOUNDARIES, J. MATER. RES., VOL 5, 1990, P 955970

21. S.P. CHEN, A.F. VOTER, R.C. ALBERS, A.M. BORING, AND P.J. HAY, THEORETICAL STUDIES OF GRAIN BOUNDARIES IN NI3AL WITH BORON OR SULFUR, SCR. METALL., VOL 23, 1989, P 217-222 22. A.F. VOTER AND S.P. CHEN, ACCURATE INTERATOMIC POTENTIALS FOR NI, AL, AND NI3AL, MATER. RES. SOC. PROC., VOL 82, 1987, P 175-180 23. S.P. CHEN, ANOMALOUS RELAXATIONS OF (0001) AND (10 1 0) SURFACES IN HCP METALS, SURF. SCI. LETT., VOL 264, 1992, P L162-L168 24. R.J. HARRISON, A.F. VOTER, AND S.P. CHEN, ATOMISTIC MODELING OF MATERIALS: BEYOND PAIR POTENTIALS, V. VITEK AND D.J. SROLOVITZ, ED., PLENUM PRESS, 1989, P 219-222 25. R.J. HARRISON, F. SPAEPEN, A.F. VOTER, AND S P. CHEN, STRUCTURE OF GRAIN BOUNDARIES IN IRON, PROC. 34TH SAGAMORE ARMY MATERIALS RESEARCH CONFERENCE (LAKE GEORGE, NY), 1990, P 651-675 (FOR FE) 26. S.P. CHEN, LOCAL VOLUME POTENTIALS FOR BCC METALS, TO BE PUBLISHED (FOR FE, NB, CR, W, AND MO) 27. S.P. CHEN, A.M. BORING, R.C. ALBER, AND P.J. HAY, THEORETICAL STUDIES OF NI3AL AND NIAL WITH IMPURITIES, HIGH TEMPERATURE ORDERED INTERMETALLIC ALLOYS III, C.C. KOCH, C.T. LIU, N.S. STOLOFF, AND A.I. TAUB, ED., PROC. MATER. RES. SOC., VOL 133, 1989, P 149-154 28. F.R. DEBOER, R. BOOM, W.C.M. MATTENS, A.R. MIEDEMA, AND A.K. NIESSEN, COHESION IN METALS: TRANSITION METAL ALLOYS, NORTH-HOLLAND, 1989 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Properties Affecting Adhesion Large impurities, such as bismuth, sulfur, phosphorus, and so on, tend to lower grain boundary adhesion, whereas small interstitial atoms, such as boron, carbon, and beryllium tend to increase cohesion in nickel, Ni3Al, and Ni3Si (Ref 1, 20, 21). Deviations from stoichiometry can either increase or decrease cohesion, depending on whether amounts of strong bonding elements, like nickel in Ni3Al, are increased or decreased. Although stoichiometry may slightly affect the ideal cohesion energy, its effect on the consumption of plastic work is potentially higher, as demonstrated in Co3Ti. Researchers (Ref 29) found that excess cobalt suppressed intergranular brittleness and rendered the material ductile, without adding strengtheners such as boron or beryllium. Sometimes impurity effects only work in off-stoichiometric compounds, such as boron in Ni3Al. Quantitative analysis of these kinds of effects have been attempted for impuritycontrolled fracture properties in iron, which has been studied extensively by Hondros and Seah (Ref 30). A simple linear formula linking cohesion and level of impurities has been extracted from many experiments (Ref 30).

References cited in this section

1. S.P. CHEN, D.J. SROLOVITZ, AND A.F. VOTER, COMPUTER SIMULATION ON SURFACES AND [001] TILT GRAIN BOUNDARIES IN NI, AL, AND NI3AL, J. MATER. RES., VOL 4, 1989, P 62-77 20. S.P. CHEN, A.F. VOTER, R.C. ALBERS, A.M. BORING, AND P.J. HAY, INVESTIGATION OF BORON EFFECT ON NI AND NI3AL GRAIN BOUNDARIES, J. MATER. RES., VOL 5, 1990, P 955970 21. S.P. CHEN, A.F. VOTER, R.C. ALBERS, A.M. BORING, AND P.J. HAY, THEORETICAL STUDIES OF GRAIN BOUNDARIES IN NI3AL WITH BORON OR SULFUR, SCR. METALL., VOL 23, 1989, P 217-222 29. J.E. HACK, D.J. STROLOVITZ, AND S.P. CHEN, A MODEL FOR THE FRACTURE BEHAVIOR OF POLYCRYSTALLINE NI3AL, SCR. METALL., VOL 20, 1986, P 1699-1704

30. E.D. HONDROS AND M.P. SEAH, INTERFACIAL AND SURFACE MICROCHEMISTRY, PHYSICAL METALLURGY, 3RD ED., R.W. CAHN AND P. HAASEN, ED., ELSEVIER SCIENCE PUBLISHER, 1983, P 855-931 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Adhesion Measurement Sikorski (Ref 31) describes two modifications of the twist-compression bonding method for adhesion testing and concludes that the measured coefficient of friction supports the adhesion theory of friction. This correlation implies that a large coefficient of adhesion (ratio of the force necessary to break a bond to the normal loading force used to produce the bond) corresponds to a large coefficient of friction. Experimentally, he showed that the coefficient of adhesion was inversely proportional to the melting point, which meant that higher melting point materials had lower adhesion. This agress with the observation that higher melting point materials have higher hardness, which also corresponds to lower adhesion coefficients. Gerkema and Miedema (Ref 32) considered lubrication of iron substrates by thin films of lead and silver with copper, platinum, or molybdenum additions by calculating surface adhesion using heats of solution and surface energies. Their predictions agree well with sliding wear experiments. Pashley and Tabor (Ref 33) measured adhesion between tungsten and nickel using techniques similar to those used in atomic force microscopy (AFM). They were able to verify surface cleanliness and ensure that the adhesion measured was between clean tungsten and nickel by using Auger spectroscopy (AES). In a subsequent publication, Pashley and Pethia again measured the surface energy of adhesion between tungsten and nickel (Ref 34). This work was done between a tungsten tip and a nickel flat in high vacuum and showed that surface forces alone could cause plastic flow. The loading/unloading curves provide information on the elastic/plastic response of the material. This in turn determines the interpretation of the adhesion. Johnson and Keller (Ref 35) used a torsion balance with a 3 mm (0.12 in.) diameter ball against a flat plate to measure the force of adhesion between silver-silver, copper-nickel, and silver-nickel couples. They concluded that mutual solubility did not affect the tendency for adhesion. However, surface contaminants did have a significant effect on the adhesion (joint strength) between silver-silver couples when the vacuum varied. The difference in joint strength for a clean surface at 130 μPa (5 × 10-10 torr) and for the same surface after a few microns of argon was 50 GPa (5 × 108 dynes/cm2). Conrad and Rice (Ref 36) studied the cohesion of fcc metals (silver, aluminum, copper, and nickel) by fracturing the metals in high vacuum, rewelding the metals by compressing the fractured surfaces, and retesting the metals by fracturing. They found that the cohesive strength of the weld increased with compressive load when both were normalized to the initial fracture strength. Results for all materials fell on a single curve. They proposed that the essentially constant cohesion coefficient was due to the fact that the contact area produced by a given compressive load was inversely proportional to the fracture stress of the metal, whereas the cohesive strength was directly proportional to the fracture stress. Acoustic emission was used to measure and monitor cohesion. In one case, this technique was used to signal the onset of fiber debonding and to determine crack initiation locations in fiber-reinforced ceramic-matrix composites (Ref 37). This approach is a valuable and viable technique for monitoring debonding, but cannot be used to provide quantitative information on either bond strength or fundamental material parameters. It can be used to locate the origin of the debonding, but the error is on the order of the dimensions used for continuum calculations, which are much larger than the atomic dimensions. In another study (Ref 38), acoustic measurements were used to monitor fatigue cracks and interfaces by defining a spring stiffness (for continuous and discontinuous spring models) that represented the degree of contact across an interface. A measured acoustic reflection coefficient was found to be in good agreement with one predicted for the same geometry. In addition, a single stress-intensity factor was found to determine bond strength.

In a similar vein, Neid (Ref 39) developed a technique for measuring bond integrity by propagating a planar shock were through the bond region. Depending on the shock energy, defects developed in the bond. The change in density was used to detect defects created at the interface, because other nondestructive test (NDT) methods were not sensitive enough to detect the small defects created. The defects were subsequently verified using optical microscopy. Itou (Ref 40) used electrical resistance to monitor crack tip contact in order to verify crack tip shielding in aluminum, steel, and cork. The results were inconclusive, although he found that there was no oscillating stress singularity around the crack tip. From this discussion, one can see that no generally quick, accurate, and easy-to-use technique has been developed to measure adhesion.

References cited in this section

31. M.E. SIKORSKI, THE ADHESION OF METALS AND FACTORS THAT INFLUENCE IT, WEAR, VOL 7, 1964, P 144-162 32. J. GERKEMA AND A.R. MIEDEMA, ADHESION BETWEEN SOLID METALS: OBSERVATIONS OF INTERFACIAL SEGREGATION EFFECTS IN METAL FILM LUBRICATION EXPERIMENTS, SURF. SCI., VOL 124, 1983, P 351-371 33. M.D. PASHLEY AND D. TABOR, ADHESION AND DEFORMATION PROPERTIES OF CLEAN AND CHARACTERIZED METAL MICRO-CONTACTS, VACUUM, VOL 10-12, 1981, P 619-623 34. M.D. PASHLEY AND J.B. PETHICA, THE ROLE OF SURFACE FORCES IN METAL-METAL CONTACTS, J. VAC. SCI. TECHNOL. A, VOL 3 (NO. 3), MAY/JUNE 1985, P 757-761 35. K.I. JOHNSON AND D.V. KELLER, JR., EFFECT OF CONTAMINATION ON THE ADHESION OF METALLIC COUPLES IN ULTRA-HIGH VACUUM, J. APPL. PHYS., VOL 38 (NO. 4), JAN 1967, P 1896-1904 36. H. CONRAD AND L. RICE, THE COHESION OF PREVIOUSLY FRACTURED FCC METALS IN ULTRA-HIGH VACUUM, METALL. TRANS., VOL 1, NOV 1970, P 3019-3029 37. T. KISHI, M. ENOKI, AND H. TSUDA, INTERFACE AND STRENGTH IN CERAMIC MATRIX COMPOSITES, MATER. SCI. ENG., VOL A143, 1991, P 103-110 38. O. BUCK, R.B. THOMPSON, D.K. REHBEIN, D.D. PALMER, AND L.J.H. BRASCHE, CONTACTING SURFACES: A PROBLEM IN FATIGUE AND DIFFUSION BONDING, METALL. TRANS. A, VOL 20, APRIL 1989, P 627-635 39. H.A. NIED, "DETERMINATION OF THE STRUCTURAL INTEGRITY OF METALLURGICALLY BONDED PLATES BY HIGH ENERGY IMPULSE TEST," SIXTH INT. CONF. ON EXP. STRESS ANALYSIS (MUNICH), SEPT 1978 40. S. ITOU, AN EXPERIMENT ON THE BEHAVIOUR AROUND THE TIPS OF AN INTERFACE CRACK, ENG. FRACT. MECH., VOL 37 (NO. 1), 1990, P 145-150 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Interfacial Analysis Electron spectroscopies are extremely valuable for the quantitative and qualitative analysis of surfaces. Fundamental information about the atomic forces and energies of the regions being probed, as well as their geometry, can be obtained. The shortcoming of these techniques in analyzing interfaces is that the interfacial volume is a small fraction of the probed volume and the effects of the interface cannot be separated from the signal. This condition can be overcome by using scanning tunneling microscopy (STM) or AFM. However, preparation of the surface is extremely difficult, the equipment requires a skilled and trained operator, the data are fundamental, and engineering information is not obtained. The crux of the problem is that there is no simple connection between condensed matter theory, micromechanics, and continuum

mechanics. The following discussion briefly reviews some of the available technical literature that describes the use of these spectroscopies and their application to bimaterial systems. Photoelectron spectroscopy has been used to study electron core-level shifts for ytterbium, aluminum, and silicon grown on molybdenum (110) (Ref 41, 42). The core-level shifts between the film and the substrate are expressed in terms of adhesion energy. In this case, the adhesion energy is the difference between the energy of the neutral and excited atoms per atom, which is readily found from the spectra. The measured adhesion energy is in good agreement with the adhesion energy determined using the Miedema technique (Ref 28). High-resolution electron microscopy (HREM), electron energy loss spectroscopy (EELS), microspectroscopy, and microdiffraction have been used to analyzed the chemical distributions on or near interface planes (Ref 43). This information can be used to support models requiring detailed stoichiometry, but does not yield adhesion values. Gerberich et al. (Ref 44) used selected-area-electron channeling patterns (SACP) and stereoimaging to quantify the strain in grains. The work was carried out under plane strain and plane stress conditions. The researchers determined strain distributions within 1 μm (40 μin.) of a stationary or growing crack in Fe-3Si single crystals with a {100} crack. Although this arrangement is not a bimaterial interface, the results reflect the effort needed to obtain fundamental materials information. Optical techniques, such as reflection (Ref 45), have also been used to probe interfaces and to aid in determining interfacial electronic structure. The data, in turn, are related to adhesion and segregation. For example, Lanning et al. (Ref 46) recently correlated reflectance spectra with surface free energy. This was then used as a measure of surface energy and the work of decohesion was calculated. The most promising technology for understanding interfaces is probably STM and associated techniques. Advances in the science and technology of scanning tunneling microscopy/spectroscopy (STM/S), AFM, photon scanning tunneling microscopy (PSTM), scanning chemical potential microscopy (SCPM), and related spectroscopies/microscopies offer new insight into interfaces (Ref 47). Low-energy electron diffraction (LEED), LEED-spot profile analysis (SPA-LEED), scanning low-energy electron microscopy (SLEEM), and photoemission electron microscopy (PEEM) are also being applied to surface analysis and adhesion. These technologies, although currently applied to the vacuum-surface interface, are appropriate tools for the study of grain boundaries and dissimilar metal surfaces. They provide excellent methods for measuring the electronic behavior at interfaces and, thus, the fundamental interactions that occur at these interfaces. This information must then be correlated to engineering properties predicted by micromechanical and continuum mechanics methods, as well as experiments.

References cited in this section

28. F.R. DEBOER, R. BOOM, W.C.M. MATTENS, A.R. MIEDEMA, AND A.K. NIESSEN, COHESION IN METALS: TRANSITION METAL ALLOYS, NORTH-HOLLAND, 1989 41. N. MARTENSSON, A. STENBORG, O. BJORNEHOLM, A. NILSSON, AND J.N. ANDERSON, QUANTITATIVE STUDIES OF METAL-METAL ADHESION AND INTERFACE SEGREGATION ENERGIES USING PHOTOELECTRON SPECTROSCOPY, PHYS. REV. LETT., VOL 60 (NO. 17), APRIL 1988, P 1731-1734 42. J.N. ANDERSON, O. BJORNEHOLM, A. STENBORG, A. NILSSON, C. WIGREN, AND N. MARTENSSON, MEASUREMENT OF METAL-METAL ADHESION AND INTERFACE SEGREGATION ENERGIES BY CORE-LEVEL PHOTOELECTRON SPECTROSCOPY. AL AND SI ON MO(110), J. PHYS., CONDENS. MATTER, VOL 1, 1989, P 7309-7313 43. R.W. CARPENTER, HIGH RESOLUTION INTERFACE ANALYSIS, MATER. SCI. ENG., VOL A107, 1989, P 207-216 44. W.W. GERBERICH, D.L. DAVIDSON, AND M. KACZOROWSKI, EXPERIMENTAL AND THEORETICAL STRAIN DISTRIBUTIONS FOR STATIONARY AND GROWING CRACKS, J. MECH. PHYS. SOLIDS, VOL 38 (NO. 1), 1990 45. A. JOHNER AND P. SCHAAF, CALCULATION OF THE REFLECTION COEFFICIENTS OF INTERFACES: A SCATTERING APPROACH, PHYS. REV. B, VOL 42 (NO. 9), SEPT 1990, P 55165526

46. B.R. LANNING, T. FURTAK, AND G.R. EDWARDS, A METHOD FOR PREDICTION OF METALCERAMIC INTERFACIAL BOND ENERGIES, SUBMITTED FOR PUBLICATION IN J. MATER. RES. 47. R.J. COLTON, C.R.K. MARRIAN, AND J.A. STROSCIO, ED., PROC. FIFTH INT. CONF. SCANNING TUNNELING MICROSCOPY/SPECTROSCOPY (SAM '90) AND THE FIRST INT. CONF. NANOMETER SCALE SCIENCE AND TECHNOLOGY (NANO I) (BALTIMORE, MD), JUNE 1990, AND J. VAC. SCI. TECH. B, SECOND SERIES, VOL 9 (NO. 2), PART II, MARCH/APRIL 1991 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Welding (Bonding) An extensive review of diffusion bonding is given by Gerken and Owczarski (Ref 48). This treatise reviewed the status of theory at that time (1965) and cited numerous examples. Eleven years later, Williams and Crossland (Ref 49, 50) highlighted cold and diffusion welding. They pointed out that the major technical concerns for these processes had not significantly changed in the interim. As recently as 1986, there was concern over the knowledge base of bonding and adhesion at interfaces (Ref 51). This group felt that detailed mechanisms of adhesion were largely missing, but that firstprinciple theories were beginning to appear. A strong need for collaboration between theorists and experimentalists was indicated. Furthermore, much information was still unknown about interfaces, and the field was ripe for expansion. This situation still applies. Described below are examples of some of the proposed models for bonding. Several mechanisms have been proposed to explain the formation of solid-state welds (Ref 52, 53). Mohamed and Washburn (Ref 52) proposed a two-stage mechanism that depends on breaking the surface oxide. Their work is supported by the work of Bay. (Ref 54), who developed a theoretical model for oxide breakage and bond formation on aluminum. Kawakatsu and Kitayama (Ref 53) describe bonding as being controlled by three metallurgical factors, which have been identified as bonds between solid-solution, two-phase, or intermetallic materials. Their studies involved bond-strength measurements and microstructural examination as a function of bonding conditions for copper-nickel, copper-silver, and iron-aluminum couples. They concluded that a minimum bonding temperature occurs at about the softening temperature of one of the materials. They also found that bond strength depends on recrystallization temperature. In the case of solidsolution materials, the interfaces tend to disappear after sufficient time above the recrystallization temperature, whereas this does not occur in two-phase and intermetallic materials. Excellent treatises on surface and interfacial segregation are provided by Dowben and Miller (Ref 55) and Hondros and Seah (Ref 30). The concept of migration or segregation of one component of a system to a surface or interface is considered in detail. This work provides insight into the behavior that must be considered when determining the possibility of adhesion between two dissimilar materials. Garmong et al. (Ref 56) describe the rate-limiting step in bonding as the complete elimination of porosity from the bond line and have developed a time-prediction model to support it. The model proceeds in two stages. First, long-wavelength asperities (roughness) are eliminated by plastic flow. Second, voids are eliminated by plastic flow and vacancy diffusion. This model is similar to that proposed by Hill and Wallach (Ref 57), who consider void elimination at the bond line. In this case, voids are created by surface asperities on the mating surfaces and pressure sintering is invoked for its elimination. The effect of grain size is also considered. Cahn (Ref 58) derived interfacial energy from thermodynamic considerations. He considered how the order parameter changes surface tension, elastic energy, and excess free energy when it affects the lattice parameter of an ordered material at an antiphase boundary. The work is not only appropriate for grain boundary interfaces in ordered materials, but is also relevant for dissimilar materials. Cline (Ref 59) reviews the difference between deformation and diffusion bonding. He defines diffusion bonding as not requiring significant deformation for bond formation. With respect to deformation bonding, Cline discusses the "film" and "energy barrier" theories and reviews the factors affecting the resultant bond. With respect to diffusion bonding, he defines two stages for bonding: microscopic deformation and diffusion. He then reviews the results for bonding aluminum and titanium and shows that both can be readily deformation bonded. The diffusion bonding of aluminum was simplified

by using a silver interlayer, whereas titanium required smooth, clean surfaces and protection from atmospheric contaminants. Derby and Wallach (Ref 60, 61) developed a thorough model for diffusion welding which accounts for: • • • • • • •

SURFACE DIFFUSION FROM SURFACE SOURCES TO A NECK VOLUME DIFFUSION FROM SURFACE SOURCES TO A NECK DIFFUSION ALONG THE BOND INTERFACE FROM INTERFACIAL SOURCES TO A NECK VOLUME DIFFUSION FROM INTERFACIAL SOURCES TO A NECK POWER-LAW CREEP DEFORMING THE RIDGE (SURFACE ROUGHNESS) PLASTIC YIELDING DEFORMING THE RIDGE VAPOR PHASE MASS TRANSPORT

Mass transport rate equations were developed in terms of the process variables (time, temperature, pressure, and material properties). The model is in reasonable agreement with experiment (Ref 62, 63). Guo and Ridley (Ref 64) consider the rate-limiting step in bonding to be the removal of interfacial voids that are due to the surface roughness and have developed a model to support this position. The main difference between their model and that of Derby and Wallach (Ref 60) is the void geometry used. Guo and Ridley use a lenticular-shaped cavity and have developed a model that agrees reasonably with the experimental data.

References cited in this section

30. E.D. HONDROS AND M.P. SEAH, INTERFACIAL AND SURFACE MICROCHEMISTRY, PHYSICAL METALLURGY, 3RD ED., R.W. CAHN AND P. HAASEN, ED., ELSEVIER SCIENCE PUBLISHER, 1983, P 855-931 48. J.M. GERKEN AND W.A. OWCZARSKI, "A REVIEW OF DIFFUSION WELDING," WELDING RESEARCH COUNCIL BULLETIN NO. 109, OCT 1965 49. J.D. WILLIAMS AND B. CROSSLAND, PART 1 COLD PRESSURE WELDING, CME, MARCH 1976, P 65-67 50. J.D. WILLIAMS AND B. CROSSLAND, PART 2 DIFFUSION BONDING, CME, APRIL 1976, P 73-75 51. J.R. SMITH, PANEL REPORT ON INTERFACIAL BONDING AND ADHESION, MATER. SCI. ENG., VOL 83, 1986 52. H.A. MOHAMED AND J. WASHBURN, MECHANISM OF SOLID STATE PRESSURE WELDING, WELD. J. RES. SUPP., SEPT 1975, P 302S-310S 53. I. KAWAKATSU AND S. KITAYAMA, STUDY ON DIFFUSION BONDING OF METALS, TRANS. JIM, VOL 18, 1977, P 455-465 54. N. BAY, MECHANISMS PRODUCING METALLIC BONDS IN COLD WELDING, WELD. J. RES. SUPP., MAY 1983, P 137S-142S 55. P.A. DOWBEN AND A. MILLER, ED., SURFACE SEGREGATION PHENOMENA, CRC PRESS, 1990 56. G. GARMONG, N.E. PATON, AND A.S. ARGON, ATTAINMENT OF FULL INTERFACIAL CONTACT DURING DIFFUSION BONDING, METALL. TRANS. A, VOL 6, JUNE 1975, P 1269-1279 57. A. HILL AND E.R. WALLACH, MODELLING SOLID-STATE DIFFUSION BONDING, ACTA METALL., VOL 37 (NO. 9), 1989, P 2425-2437 58. J.W. CAHN, INTERFACIAL FREE ENERGY AND INTERFACIAL STRESS: THE CASE OF AN INTERNAL INTERFACE IN A SOLID, ACTA METALL., VOL 37 (NO. 3), 1989, P 773-776 59. C.L. CLINE, AN ANALYTICAL AND EXPERIMENTAL STUDY OF DIFFUSION BONDING, WELD. J., NOV 1966, P 481-S TO 489-S 60. D. DERBY AND E.R. WALLACH, THEORETICAL MODEL FOR DIFFUSION BONDING, MET. SCI.,

VOL 16, JAN 1982, P 49-56 61. D. DERBY AND E.R. WALLACH, DIFFUSION BONDING: DEVELOPMENT OF THEORETICAL MODEL, MET. SCI., VOL 18, SEPT 1984, P 427-431 62. D. DERBY AND E.R. WALLACH, DIFFUSION BONDS IN IRON AND A LOW-ALLOY STEEL, J. MATER. SCI., VOL 19, 1984, P 3149-3158 63. D. DERBY AND E.R. WALLACH, DIFFUSION BONDS IN COPPER, J. MATER. SCI., VOL 19, 1984, P 3140-3148 64. Z.X. GUO AND N. RIDLEY, MODELLING OF DIFFUSION BONDING OF METALS, MATER. SCI. TECH., VOL 3, NOV 1987, P 945-953 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Interface Formation When considering the compression needed to bond dissimilar materials, the deformation-mechanism map work of Frost and Ashby (Ref 65) is helpful. The maps are in a stress/temperature space in which a single deformation mechanism is dominant. The stress used is a normalized tensile stress (tensile stress/shear modulus), and the temperature is the homologous temperature T/Tm, where T is the test temperature and Tm is the melting point of the material being tested, where both temperatures are expressed in degrees K. From these maps and the corresponding constitutive equations, contour plots of constant strain rate can be plotted. At this point, knowing any two values of strain rate, stress, or temperature, a point on the map will give the third variable and thee dominant deformation mechanism. This information can be a guide to determining the bonding mechanism when considering one of the previously discussed mechanisms that requires elimination of defects at the interface. Dammer et al. (Ref 66) describe a study where type 347 stainless steel was bonded to itself with several different interlayers. The interlayers were electrolytic-grade copper foil, electrolytic tough pitch (ETP) copper sheet, 80Au-20Cu foil, 72Ag-28Cu foil, and several combinations of silver, copper, and gold electroplates. They successfully bonded the specimens at temperatures below 540 °C (1005 °F), but found that the intermediate materials were essential to forming the weld. Bachin (Ref 67) discusses the interactions of metals and oxides when diffusion welding metals to metals and to ceramics. He discusses the reactions and exchanges at the interface by considering the formation of spinels and oxides. This leads to the conditions under which each is known to be operative and to the possibility of diffusion across the resulting interface. For example, Al + SiO2 proceeds by exchange:

4AL + 3SIO2 → 2AL2O3 + 3SI

(EQ 7)

When thin silicon dioxide is joined through a thin intermediate interlayer of aluminum, new phases precipitate at the dividing boundaries and the solid solution is decomposed when the limit of solubility is reached for the reaction products. This behavior results in high internal stresses and resultant microcracks. Caution is therefore required when welding these systems. A thorough review of the thermodynamics involved for the systems being considered is essential. Munir (Ref 68) developed equations that predict the amount of time required to dissolve oxides into their host metals for systems that have unstable oxides with respect to dissolution. He considered niobium, zirconium, titanium, copper, tantalum, cobalt, nickel, and iron and found reasonable agreement with the theory. Wang and Vehoff (Ref 69) discuss the segregation of atoms to interfaces. Their work deals with grain boundaries as interfaces, and they conclude that the cohesive energy of the interface at a fixed chemical potential is always less than the cohesive energy at a fixed concentration. The concept of segregation to the interface is probably important for bimaterial interfaces because of the different chemical potentials, diffusion rates, and, thus, concentrations that can exist. Because

these parameters are all temperature dependent, the bonding process and the service temperature will affect the interface segregation and, thus, its performance. This work provides a way to consider these effects.

References cited in this section

65. H.J. FROST AND M.F. ASHBY, DEFORMATION MECHANISM MAPS, PERGAMON PRESS, 1982 66. P.A. DAMMER, R.E. MONROE, AND D.C. MARTIN, FURTHER STUDIES OF DIFFUSION BONDING BELOW 1000 °F, WELD. J., MARCH 1969, P 116-S TO 124-S 67. V.A. BACHIN, INTERACTION BETWEEN METALS AND THE OXIDES OF CERAMIC MATERIALS DURING DIFFUSION WELDING, AVT. SVARKA, NO. 10, 1982, P 15-17 68. Z.A. MUNIR, A THEORETICAL ANALYSIS OF THE STABILITY OF SURFACE OXIDES DURING DIFFUSION WELDING OF METALS, WELD. J., DEC 1983, P 333-S TO 336-S 69. J.-S. WANG AND H. VEHOFF, THE EFFECT OF THE MOBILITY OF SEGREGATED ATOMS ON INTERFACIAL EMBRITTLEMENT, SCR. METALL., VOL 25, 1991, P 1339-1344 Fundamentals of Metal and Metal-to-Ceramic Adhesion Ray Dixon and S.P. Chen, Los Alamos National Laboratory

Strength of Interfaces One of the most important and probably the most frequent questions: What is the strength of the interfacial bond? This question has often been answered by tensile testing. The difficulty is that the failure is often not at the interface and that it must be forced there. Additionally, there is concern that the strain is not one-dimensional across the bond and that the bond experiences triaxial strain in service. For these reasons, there has been considerable interests in determining fracture toughness and in detecting the failure of interfaces. Several reviews on the importance and extent of interfacial science were recently presented in one technical journal (Ref 70). Three reviews are noteworthy because they discuss the mechanical properties of interfaces (Ref 71, 72, 73). The efforts of Hutchinson and Suo (Ref 71) bring together much of the recent work on elastic fracture phenomena for bimetallic systems and builds on it to support work being done to understand mixed-mode failure in systems with interfaces. Shih (Ref 72) summarizes progress on interfacial mechanics and points out that fracture toughness, at various mode mixities, is needed to characterize fracture resistance of a weak plane in a composite. An application of this analysis has been demonstrated on the alumina/niobium metal-ceramic system by Stout et al. (Ref 74). Comninou (Ref 73) reviews the interface crack problem from the standpoint of open and closed crack tips and describes preliminary experimental results on fatigue and fracture of interface cracks in aluminum bonded by epoxy. Interface Strength Models. There has been considerable progress in analyzing the fracture behavior of interfaces with

various configurations. This work has led to the development of specimen geometries for systems with differing physical and mechanical properties. The analysis work is reviewed in this section and the experimental work is described in the section "Interface Strength Measurements" of this article. Suo and Hutchinson (Ref 75) have provided a complete solution to the semi-infinite crack lying along the interface of two infinite elastic layers under edge loads. This solution allows the determination of material properties for complex geometries, as well as the effect of welding two dissimilar materials. Jensen (Ref 76) demonstrates that shielding the full effect of mode II and III loading is important for interface cracks in brittle materials and that a proper interface fracture criteria needs to be modeled. He discusses a method for comparing different phenomenological fracture criteria and sets up micromechanical models describing shielding for interface geometries with kinks. The models of He and Hutchinson (Ref 77) explain conditions for interfacial crack propagation (kinking) into material on either side of the interface. Their models suggest that a crack propagates along an interface as long as the compliant material is tough and that the stiff material is at least as tough as the interface. They also show that the energy release rate is increased if the crack kinks into the more-compliant material and is diminished if it kinks into the stiff material.

Qu and Li (Ref 78) also analyzed cracks and dislocations at the interface between two anisotropic elastic half spaces. They obtained an explicit and real solution to the dislocation problem in the Stroh (Ref 79) formalism and showed that a planar interface between dissimilar anisotropic solids is completely characterized by no more than nine independent parameters. The work by Li and Ting (Ref 80, 81) is also relevant. Evans et al. (Ref 82, 83) concluded that fracture energy is not unique and is usually greater than the thermodynamic work of adhesion. The nonuniqueness is related to the mixity experienced by interface cracks. As a result, the fracture energy is influenced by the choice of test specimen. This result implies that when considering fracture energy for specific material combinations, the test geometry must be known, reported, and maintained. This also means that it is now possible to systematically study interfacial mechanical behavior and produce meaningful data on interface fracture toughness. Fischmeister et al. (Ref 84) modeled simple crack propagation, dislocation emission from crack tips, and twinning using atomistic calculations (molecular dynamics) embedded in a finite-element continuum. They also studied the behavior of cracks propagating in tungsten and iron on different planes and in different directions within these planes. Their results confirm that {100} should be the easiest cleavage plane in these body-centered cubic (bcc) tungsten and iron crystals. They further studied the propagation direction restrictions for cracks on particular planes. The modeling work was supported by experiments to verify the crack propagation behavior in WC-Co alloy, high-speed steel consolidated with nickel particles, and nickel coated on high-speed steel. Chen et al. (Ref 1, 9, 20, 21) studied crack growth in nickel, Ni3Al, Ni3Al plus boron, Ni3Al plus sulfur, and iridium grain boundaries using local volume potentials. They found that cohesive strength is important in determining grain boundary fracture and can be strongly affected by impurities and stoichiometry. Another approach to modeling fracture is given by Needleman (Ref 85), who used a cohesive zone-type interface model. His model takes full account of finite geometry changes and is used to study decohesion of a viscoplastic block from a rigid substrate. He found that the normal traction versus separation relation had an exponential form consistent with the universal binding energy relationship of Rose et al. (Ref 86) and of Ferrante and Smith (Ref 87). Ting (Ref 88) showed that oscillatory displacement at the interfacial crack surface in a bimaterial under a twodimensional deformation depends only on the material properties and is independent of specimen orientation. The exception is that bimaterials always produce oscillation and that there is an exception for one choice of relative orientation in one material combination. He also shows that for two different orthotropic materials, the oscillation depends on whether or not the Voigt elastic constants combinations, C11 + C22 + C12 , for the two materials are different. Cook (Ref 89) explains crack propagation threshold in brittle materials through the work done by the applied loading and that which is needed to create new surfaces during crack propagation. He showed that the mechanical energy release rate was a measure of the equilibrium surface energy of the material and the chemical environment. The latter affects the surface tension of both the fresh and aged crack. Reynolds et al. (Ref 90) investigated the difference between the concepts of cracks with zero normal and shear traction over the entire face of the crack and cracks with normal and shear traction specified at zero on the crack surface, except near the tip, where zero shear fraction and continuity of displacements normal to the crack were required. Of these two approaches, the first leads to oscillatory stress singularities and interpenetration of the material near the crack tip, whereas the second shows that a small region of contact exists at the tip, regardless, of the loading. The two methods require different treatments in terms of stress intensity, but predict the same strain-energy release rate. The work was done using the finite-element method (FEM) and, thus, also served to verify the formulation of each problem by FEM analysis. The results showed that the two formulations could be modeled by the absence or presence of interface elements that prevent interpenetration of the crack surfaces. Additionally, fracture parameters were calculated and compared well with previous analytic solutions. Interface Strength Measurements. Based on the theoretical analysis described above, as well as their own models,

the authors identified below have carried out excellent experimental work in measuring the mechanical properties of interfaces. Charalambides et al. (Ref 91) developed a test procedure to measure interface fracture resistance under mixed-mode conditions. Strain-energy release rates and stress-intensity factors can be obtained using this test. O'Dowd et al. (Ref 92) developed FEM model solutions and tested several geometries to determine interfacial fracture toughness as a function of mixity. They were able to explain near-tip contact and variation of mixity with distance. They also developed a procedure

for determining the effect of residual stress on the stress-intensity factor. Brady et al. (Ref 93) developed a buckled plate test to determine the transverse fracture toughness of unidirectional continuous-fiber composites. Their test was found to be independent of sample geometry and testing conditions. Mecholsky and Baker (Ref 94) developed a double cantilever beam specimen with a chevron notch to measure fracture toughness of ceramic-metal interfaces. The specimen is easy to make and introduces its own crack, and the test is easy to perform. Zhu and Byrne (Ref 95) looked at the effect of interfacial defects by bonding and subsequent tensile, fatigue, and photomicroscopy tests. They measured void ratios at the bond line and found the defects smaller than those detectable by current nondestructive test methods (~0.2 to 0.5 mm2 by 1 to 3 μm, or 3 × 10-4 to 8 × 10-4 in.2, or 40 to 120 μin., thick) could still significantly affect bond strength. They concluded that fatigue testing is a better test method than tensile testing because it is sensitive to a void ratio of approximately 27%, whereas tensile tests were not. Thouless (Ref 96) used epoxy-bonded glass slides to investigate interfacial failure under mixed-mode loading. He found that when the scale is sufficiently large, the sample could be considered homogeneous with a weak planar interface. This result produced a correlation between apparent fracture resistance and the degree of mixity. At a very detailed scale, the lack of homogeneity required that the crack-tip stress be determined by the position of the crack within the interface (epoxy). This condition leads to a change in failure mechanisms, whereby under pure mode-I conditions, the crack propagates entirely within the interfacial layer, whereas mode-II effects drive the crack toward the glass-epoxy interface. Nickel specimens with known surface roughnesses were diffusion welded under controlled conditions of temperature and pressure (Ref 97). The welds were interrupted at several different time intervals and the resulting interface was examined. The defect sizes were successfully related to surface tension and diffusion. It was also found that temperature increases produced a greater effect on reducing interfacial defects than did pressure. Meriin and Sliozberg (Ref 98) make the point that sound welds (bonds) cannot be obtained unless the voids are eliminated at the interface, which requires high temperatures and pressures. It further requires that mechanical tests be used as the principal detection method. This article has shown that there has been and continues to be active interest in interfacial behavior. The science is sufficiently complex that numerous approaches are useful to evaluate interfacial behavior, depending on the specific goal. However, it is obvious that the seemingly unrelated approaches all seek to understand the interface. It is therefore incumbent upon scientists in this field to be familiar with all developments concerning adhesion so that they can contribute to its understanding and future advancement.

References cited in this section

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83. A.G. EVANS, M. RUHLE, B.J. DALGLEISH, AND P.G. CHARALAMBIDES, THE FRACTURE ENERGY OF BIMATERIAL INTERFACES, METALL. TRANS. A, VOL 21, SEPT 1990, P 2419-2429 84. H.F. FISCHMEISTER, H.E. EXNER, M.-H. POECH, S. KOHLHOFF, P. GUMBSCH, S. SCHMAUDER, L.S. SIGL, AND R. SPEIGLER, MODELING FRACTURE PROCESSES IN METALS AND COMPOSITE MATERIALS, Z. METALLKDE, BD 80, 1989, P 839-846 85. A. NEEDLEMAN, AN ANALYSIS OF TENSILE DECOHESION ALONG AN INTERFACE, J. MECH. PHYS. SOLIDS, VOL 38 (NO. 3), 1990, P 289-324 86. J.H. ROSE, J. FERRANTE, AND J.R. SMITH, UNIVERSAL BINDING ENERGY CURVES FOR METALS AND BIMETALLIC INTERFACES, PHYS. REV. LETT., VOL 47 (NO. 9), AUG 1981, P 675678 87. J. FERRANTE AND J.R. SMITH, THEORY OF THE BIMETALLIC INTERFACE, PHYS. REV. B, VOL 31 (NO. 6), MARCH 1985, P 3427-3434 88. T.C.T. TING, INTERFACE CRACKS IN ANISOTROPIC BIMATERIALS, J. MECH. PHYS. SOLIDS, VOL 38 (NO. 4), 1990, P 505-513 89. R.F. COOK, CRACK PROPAGATION THRESHOLDS: A MEASURE OF SURFACE ENERGY, J. MATER. RES., VOL 1 (NO. 6), NOV/DEC 1986, P 852-860 90. R.R. REYNOLDS, K. KOKINI, AND G. CHEN, THE MECHANICS OF THE INTERFACE CRACK USING THE FINITE ELEMENT METHOD, J. ENG. MATER. TECHNOL., VOL 112, JAN 1990, P 38-43 91. P.G. CHARALAMBIDES, J. LUND, A.G. EVANS, AND R.M. MCMEEKING, "A TEST FOR DETERMINING THE FRACTURE RESISTANCE OF BIMATERIAL INTERFACES," REPORT M87-4, MATERIALS DEPARTMENT, COLLEGE OF ENGINEERING, UC SANTA BARBARA 92. N.P. O'DOWD, C.F. SHIH, AND M.G. STOUT, TEST GEOMETRIES FOR MEASURING INTERFACIAL FRACTURE TOUGHNESS, INT. J. SOLIDS AND STRUCT., VOL 29, 1992, P 571-589 93. R.L. BRADY, R.S. PORTER, AND J.A. DONOVAN, A BUCKLED PLATE TEST TO MEASURE INTERFACIAL TOUGHNESS IN COMPOSITES, J. MATER. SCI., VOL 24, 1989, P 4138-4143 94. J.J. MECHOLSKY AND L.M. BAKER, A CHEVRON-NOTCHED SPECIMEN FOR FRACTURE TOUGHNESS MEASUREMENTS OF CERAMIC-METAL INTERFACES, CHEVRON-NOTCHED SPECIMENS: TESTING AND STRESS ANALYSIS, STP 855, J.H. UNDERWOOD, S.W. FREIMAN, AND F.I. BARATTA, ED., ASTM, 1984, P 324-336 95. X. ZHU AND J.G. BYRNE, EFFECT OF INTERFACE DEFECTS ON THE MECHANICAL PROPERTIES OF AUTOGENOUS DIFFUSION BONDED JOINTS, J. MATER. ENG., VOL 13, 1991, P 207-211 96. M.D. THOULESS, FRACTURE OF A MODEL INTERFACE UNDER MIXED-MODE LOADING, ACTA METALL. MATER., VOL 38 (NO. 6), 1990, P 1135-1140 97. V.S. GOSTOMEL'SKII, E.S. KARAKOZOV, AND A.P. TERNOVSKII, THE IMPORTANCE OF DIFFUSION AND SURFACE TENSION IN THE FORMATION OF CONTACT DURING DIFFUSION WELDING, AVT. SVARKA, NO. 4, 1980, P 28-31 98. A.V. MERIIN AND S.K. SLIOZBERG, SOME CHARACTERISTICS OF JOINTS IN THE SOLIDPHASE WELDING OF METALS, SVAR. PROIZ., NO. 5, 1973, P 22-23 Fundamentals of Friction Welding J.W. Elmer and D.D. Kautz, Lawrence Livermore National Laboratory

Introduction FRICTION WELDING (FRW) is a solid-state welding process in which the heat for welding is produced by the relative motion of the two interfaces being joined. This method relies on the direct conversion of mechanical energy to thermal energy to form the weld, without the application of heat from any other source. Under normal conditions no melting

occurs at the interface. Figure 1 shows a typical friction weld, in which a nonrotating workpiece is held in contract with a rotating workpiece under constant or gradually increasing pressure until the interface reaches the welding temperature. The rotational speed, the axial pressure, and the welding time are the principal variables that are controlled in order to provide the necessary combination of heat and pressure to form the weld. These parameters are adjusted so that the interface is heated into the plastic temperature range where welding can take place. Once the interface is heated, axial pressure is used to bring the weld interfaces into intimate contact. During this last stage of the welding process, atomic diffusion occurs while the interfaces are in contact, allowing a metallurgical bond to form between the two materials.

FIG. 1 SCHEMATIC SHOWING FUNDAMENTAL STEPS IN THE FRICTION WELDING PROCESS. (A) ONE WORKPIECE IS ROTATED, AND THE OTHER WORKPIECE IS HELD STATIONARY. (B) BOTH WORKPIECES ARE BROUGHT TOGETHER, AND AXIAL FORCE IS APPLIED TO BEGIN THE UPSETTING PROCESS. (C) WORKPIECE ROTATION IS STOPPED, AND THE UPSETTING PROCESS IS COMPLETED

Friction welding involves heat generation through friction abrasion (Ref 1, 2, 3), heat dissipation, plastic deformation, and chemical interdiffusion. The interrelation among these factors during FRW to complications when trying to develop predictive models of the friction welding process. However, from a qualitative standpoint the process is well understood through empirical FRW studies that have been performed on a wide variety of materials. Five qualitative factors influence the quality of a friction weld (Ref 4): • • • • •

RELATIVE VELOCITY OF THE SURFACES APPLIED PRESSURE SURFACE TEMPERATURE BULK MATERIAL PROPERTIES SURFACE CONDITION AND PRESENCE OF SURFACE FILMS

The first three factors are related to FRW, while the last two are related to the properties of the materials being joined. During FRW, the relative velocity, the applied pressure, and the duration of the force are the three variables that are controlled. The effect of these variables on weld quality will be discussed for the two basic friction welding processes: direct-drive welding and inertia-drive welding. The surface temperature is the critical parameter for ensuring good welds and is dependent on the processing conditions and the materials being joined. Although the surface temperature is not measured or directly controlled the effects of insufficient or excessive temperature are generally apparent through visual examination of the finished weld. The bulk material properties and the condition of the surfaces being joined affect both the frictional forces and the forging characteristics of the materials being joined. These factors will be discussed for the friction welding of both similar-material and dissimilar-material combinations.

Commercial FRW applications employ a number of variations on the basic FRW concepts. These variations were developed to accommodate different part geometries and to produce different metallurgical effects (Ref 5) and will not be discussed in this article.

References

1. T.H. HAZLET, FUNDAMENTALS OF FRICTION WELDING, SOURCE BOOK ON INNOVATIVE WELDING PROCESSES, AMERICAN SOCIETY FOR METALS, 1981, P 11-36 2. F.P. BOWDEN AND D. TABOR, THE FRICTION AND LUBRICATION OF SOLIDS, PART I, OXFORD UNIVERSITY PRESS, 1954 3. J. GODDARD AND H. WILMAN, A THEORY OF FRICTION AND WEAR DURING THE ABRASION OF METALS, WEAR, VOL 5, 1962 4. V.I. VILL, FRICTION WELDING OF METALS, TRANSLATED FROM THE RUSSIAN, AMERICAN WELDING SOCIETY, 1962 5. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 739-763 Fundamentals of Friction Welding J.W. Elmer and D.D. Kautz, Lawrence Livermore National Laboratory

Friction Welding Technology There are two principal FRW methods: direct-drive welding and inertia-drive welding. Direct-drive FRW, sometimes called conventional friction welding, uses a motor running at constant speed to input energy to the weld. Inertia-drive friction welding, sometimes called flywheel friction welding, uses the energy stored in a flywheel to input energy to the weld. These two FRW technologies produce inherently different metallurgical effects at the joint interface. Both FRW technologies can be applied through different types of relative motion in order to generate the friction necessary to form the weld. The most common FRW geometry is that shown in Fig. 1, in which one cylindrical component is held stationary and the other is rotated. However, in other methods, both components are rotated in opposite directions, or two-stationary components are pushed against a rotating piece positioned between them. Additional forms of FRW, such as radial, orbital, and linear reciprocating motions, have been developed for special part geometries. These alternate methods are discussed elsewhere (Ref 5, 6). Direct Drive Welding In direct-drive FRW, a machine resembling a lathe is equipped with a brake and clutch, a means of applying and controlling axial pressure, and a weld-cycle timer and displacement controller. The operation of a direct-drive machine consists of a friction phase where heat is generated, a stopping phase where the rotation is terminated, and a forging phase where the pressure is applied to join the pieces. The relationships among the process variables are shown in Fig. 2, which plots the rotational speed and the axial pressure as a function of time for typical weld. The time required to stop the spindle is also an important variable because it affects the weld temperature and the timing of the forging force.

FIG. 2 PLOT OF SELECTED PARAMETERS VERSUS TIME RELATIVE TO THE THREE PHASES OF THE DIRECTDRIVE FRW PROCESS

The forging phase starts at the instant when higher pressure (that is, a larger forging force) is applied in the weld cycle. Thus, the forging phase actually starts somewhere in the stopping phase. In general, the larger forging force can be applied (case 1) while the spindle is decelerating in the stopping phase (Fig. 2) or (case 2) after the spindle has stopped rotating at the end of the stopping phase. The difference in the two applications of the larger forging force is the presence of a second friction peak. In case 1, the torque will rise again to reach a second peak before dropping. This produces a torsional force. In case 2, especially when the stopping phase is very short due to rapid braking, frictional torque does not rise but actually starts to decrease at the onset of the forging phase. In this case, there is no torsional force, and forging is affected only by the upsetting force. The speed of rotation is the least sensitive process variable in that it can be varied over a wide range if heating time and pressure are properly controlled. For steels, the recommended peripheral velocity varies from 75 to 215 m/min (250 to 700 ft/min) (Ref 7). In general, higher speeds correspond to low weld heat inputs and are used to weld heat-sensitive materials such as hardenable steels. The friction force is generally applied gradually to the weld to help overcome the initial contact-torque peak. For carbon steels, a friction pressure of about 70 MPa (10 ksi) at the interface area is required to form a good joint. After the drive motor is disengaged from the workpiece, the forging force is applied to complete the weld. Typical forging forces for carbon steel are of 140 MPa (20 ksi) at the weld interface. Inertia-Drive Welding The inertia-drive FRW method uses a similar type of machine except that the spindle holding the rotating piece is attached to a flywheel. The flywheel controls the energy input to the weld. The moment of inertia of the flywheel is an important variable that is adjusted by adding or removing flywheels. The amount of energy stored in the flywheel is controlled by its speed. Once the spindle is at the correct speed, the drive system is disengaged, leaving a rotating flywheel mass. Axial pressure is then applied and held constant throughout the welding process. The applied pressure results in a decrease in the rotational speed, typically referred to as deceleration. In some cases, when the spindle has either nearly stopped or come to a complete stop, a higher forging force may be used. Figure 3 illustrates the inertia-drive FRW process, which is similar to the direct-drive method in that the weld typically takes place in two stages: friction and forging. However, some weld schedules do not require a forging stage.

FIG. 3 PLOT OF SELECTED PARAMETERS VERSUS TIME RELATIVE TO THE TWO PHASES OF THE INERTIADRIVE FRW PROCESS

The major difference between the direct-drive and inertia-drive methods is the friction speed. In inertia welding, the friction speed continuously decreases during the friction stage, while in direct welding the friction speed stays constant. The heat generated by the plastic deformation of materials at the faying surfaces, not the heat generated by friction in the friction phase, is the primary source in the forging phase of preventing rapid decrease of the temperature at the interface. The process variables that control the characteristics of an inertia weld are the flywheel size (moment of inertia), the flywheel speed, and the axial pressure. The weld energy is related to the first two variables and is a fixed quantity once they have been determined. The kinetic energy in the flywheel at any time during the weld is given by: E=

S ².1 C

where E is the energy (ft · lbf, or J), I is the moment of inertia (lb/ft2, or kg/m2), S is the flywheel speed (rev/min), and C is a conversion constant that is equal to 5873 for English units or 182.4 for metric units. The constant C is derived from: E=

1 mv ² 2

where v = ωr. Because ω= 2πs, v = 2πsr: 1 m(2π )² s ² r ² 2 1 E = mr ²(2π )² s ² 2 E=

Making mr2 = I, the previous equation becomes: E=

1 l (2π )² s ² 2

In SI units, the previous equation becomes:  m².Kg  1  1   min  E= = (2π )² s ²  l [ m².Kg ]      s²  2  min ²   60 s  1 (2π )² E= s ²l 2 60² Is ² E= C

2

Solving for C in the previous equation: C=

2*60² 60² = = 182.378 (2π ) 2 2π 2

The energy stored in the flywheel is proportional to its speed of rotation squared, S2. Therefore, a wide range of energy levels can be obtained without changing the flywheel to accommodate changes in part geometries. For large changes in the parts being joined, the capacity of an inertia welding machine can be modified by changing the flywheel moment of inertia if necessary. Flywheel Energy. The moment of inertia of the flywheel is selected to produce both the desired amount of kinetic

energy and the desired amount of forging. Forging results from the characteristic increase in torque that occurs at the weld interface as the flywheel slows and comes to rest. This increased torque, in combination with the axial pressure, produces forging. Because forging begins at some critical velocity, the amount of forging depends on the amount of energy remaining in the flywheel, which is a linear function of the flywheel moment of inertia. Large, low-speed flywheels produce greater forging force than small, high-speed flywheels even though they contain the same amount of kinetic energy. Although small, medium, and large amounts of flywheel energy produce similar heating patterns, the amount of energy greatly affects the size and shape of the weld upset, as shown in Fig. 4(a) for similar-metal joints.

FIG. 4 SCHEMATIC SHOWING EFFECT OF WELDING PARAMETERS ON THE FINISHED WELD NUGGET OBTAINED WHEN SIMILAR METALS ARE WELDED USING INERTIA-DRIVE FRW EQUIPMENT. (A) FLYWHEEL ENERGY. (B) INITIAL PERIPHERAL VELOCITY OF WORKPIECE. (C) AXIAL PRESSURE. SOURCE: REF 7

Peripheral Velocities. There is an optimum range of peripheral or linear surface velocities for each combination of

metals being joined. For welding steel to steel, the recommended initial peripheral velocity ranges from 90 to 460 m/min (300 to 1500 ft/min). However, welds can be made at velocities as low as 85 m/min (275 ft/min). Figure 4(b) shows the effect of initial peripheral velocity on weld shape for similar-metal welds. Axial Pressure Versus Peripheral Velocities. The effect of varying the axial pressure is opposite to the effect of

varying the velocity. Welds made at low axial pressure resemble welds made at medium velocity relative to the formation of weld upset and heat-affected zones (HAZ), as shown in Fig. 4(c). Excessive pressure produces a weld that has poor quality at the center and has a large amount of weld upset, similar to a weld made at a low velocity.

References cited in this section

5. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 739-763 6. RECOMMENDED PRACTICES FOR FRICTION WELDING, ANSI/AWS C6. 1-89, AMERICAN WELDING SOCIETY, 1989

7. METALS HANDBOOK, 9TH ED., VOL 6, AMERICAN SOCIETY FOR METALS, 1983 Fundamentals of Friction Welding J.W. Elmer and D.D. Kautz, Lawrence Livermore National Laboratory

Metallurgical Parameters Friction welding can be used to join a wide range of similar and dissimilar materials. Metals, ceramics, metal-matrix composites (MMC), and polymers have all been joined by FRW, and many of the dissimilar-metal combinations that cannot be joined by conventional fusion welding techniques are readily joined by FRW methods. This section summarizes some of the metals that have been joined by FRW and discusses the metallurgical considerations that govern the properties of the resulting weld. Joining of Similar Metals The two general requirements for forming good friction welds are, first, that the materials to be joined can be forged and, second, that the materials can generate friction at the weld interface. The first requirement eliminates similar-metal welds in brittle materials such as ceramics, cast irons, and cemented carbides. However, ductile materials can sometimes be joined to these materials. The second requirements eliminates materials that contain alloying that provide dry lubrication. Free-machining additives to steel, graphite-containing alloys such as cast iron, and lead alloys may suffer from this requirement. Almost all other metal alloys can be welded to themselves by FRW techniques. Table 1 summarizes a number of common similar-metal weld joints that have been made using inertia-drive FRW. Metallic alloys known to form high-quality FRW joints include alloys based on aluminum, copper (copper-nickel, brass, bronze), iron (low-alloy steel, tool steel, stainless steel, maraging steel), nickel, titanium, tantalum, and many others (Ref 5, 6). Near full-strength metallurgical bonds can be produced for a very wide range of similar-metal alloy friction welds. The microstructure and mechanical properties of inertia-welded similar-metal joints for the following alloys can be found in the sources listed:

TABLE 1 PARAMETERS FOR INERTIA-DRIVE FRICTION WELDING OF TWO 25 MM (1 IN.) DIAM BARS MADE OF SIMILAR METALS WORK METAL

SPINDLE SPEED, REV/MIN

AXIAL FORCE KN LBF × 103

FLYWHEEL SIZE(A) KG · LB · M2 FT2

WELD ENERGY KJ FT · LBF × 103

METAL LOSS(B) MM IN.

TOTAL TIME(C), S

1018 STEEL 1045 STEEL 4140 STEEL INCONEL 718 MARAGING STEEL TYPE 410 STAINLESS TYPE 302 STAINLESS COPPER (COMMERCIALLY PURE) COPPER ALLOY 260 (CARTRIDGE BRASS, 70%) TITANIUM ALLOY, TI6AL-4V ALUMINUM ALLOY 1100

4600 4600 4600 1500 3000 3000 3500 8000

53 62 67 220 90 80 80 22

12 14 15 50 20 18 18 5

0.28 0.33 0.35 5.48 0.84 0.84 0.59 0.04

6.7 7.8 8.3 130.0 20.0 20.0 14.0 1.0

33 38 41 68 41 41 41 14

24 28 30 50 30 30 30 10

2.5 2.5 2.5 3.8 2.5 2.5 2.5 3.8

0.10 0.10 0.10 0.15 0.10 0.10 0.10 0.15

2.0 2.0 2.0 3.0 2.5 2.5 2.5 0.5

7000

22

5

0.05

1.2

14 10

3.8

0.15

0.7

6000

36

8

0.07

1.7

22 16

2.5

0.10

2.0

5700

27

6

0.11

2.7

20 15

3.8

0.15

1.0

ALUMINUM ALLOY 6061

5700

31

7

0.13

3.0

23 17

3.8

0.15

1.0

Source: Ref 7

(A) MOMENT OF INERTIA OF THE FLYWHEEL. (B) TOTAL SHORTENING OF THE WORKPIECES DURING WELDING. (C) SUM OF HEATING TIME PLUS WELDING TIME. The relative ease of friction welding metals to themselves is related to the matching properties at the weld interface. Because the materials properties are matched, heat is distributed uniformly on both sides of the joint, and the deformation characteristics are identical on both sides of the joint. This results in symmetric welds with good properties. In general, the process variables do not vary significantly for different alloys within a given class of materials. However, there can be a significant variation in processing variables between different classes of materials (Table 1). Because FRW generates localized heating at the interface, the HAZ is subject to rapid cooling due to heat transfer to the cold base metal. This rapid quenching may sufficiently alter the mechanical properties of the base metal in the HAZ region to require postweld heat treatment. For example, in order to restore ductility, stress relieving or tempering may be required to friction weld steels with hardenability greater than that of AISI 1035 (Ref 5, 6). In addition, age-hardenable alloys will lose strength in the HAZ during welding and may require postweld solution heat treating and/or postweld aging to restore their strength. Other alloys, such as those that obtain their strength from cold working, will lose strength in the HAZ of the weld, and their properties cannot be restored with postweld treatments. Joining of Dissimilar Metals While many similar-metal FRW joints are produced because of economic considerations, many dissimilar-metal FRW joints are produced because there are no alternative welding methods that can be used. Examples of these types of joints include dissimilar-metal combinations with widely different melting points and dissimilar-metal combinations that form incompatible phases when fusion welded. Table 2 gives parameters used for inertia welding several common dissimilarmaterial combinations.

TABLE 2 PARAMETERS FOR INERTIA-DRIVE FRICTION WELDING OF TWO 25-MM (1-IN.) DIAM BARS MADE OF DISSIMILAR METALS WORK METAL

SPINDLE SPEED, REV/MIN

AXIAL FORCE KN LBF × 103

FLYWHEEL SIZE(A) KG · LB · M2 FT2

WELD ENERGY KJ FT · LBF × 103

METAL LOSS(B) MM IN.

TOTAL TIME(C), S

Copper to 1018 steel M2 tool steel to 1045 steel Nickel alloy 718 to 1045 steel Type 302 stainless to 1020 steel Sintered high-carbon steel to 1018 Aluminum 6061 to type 302 stainless

8000 3000

22 180

5 40

0.06 1.14

1.4 27.0

20 54

15 40

3.8 2.5

0.15 0.10

1.0 3.0

1500

180

40

5.48

130.0

68

50

3.8

0.15

2.5

3000

80

18

0.84

20.0

41

30

2.5

0.10

2.5

4600

53

12

0.35

8.3

41

30

2.5

0.10

2.5

5500

22 5 (D) 67 15(D)

0.16 ...

3.9 ...

5.1 ...

0.20 ...

3.0 ...

Copper to aluminum alloy 1100

2000

33

0.46

11.0

27 20 . . ... . 10 7.5

5.1

0.20

1.0

7.5

(A) MOMENT OF INERTIA OF THE FLYWHEEL. (B) TOTAL SHORTENING OF THE WORKPIECES DURING WELDING.

(C) SUM OF HEATING TIME PLUS WELDING TIME. (D) THE 22 KN (5000 LBF) FORCE IS APPLIED DURING THE HEATING STAGE OF THE WELD; FORCE IS SUBSEQUENTLY INCREASED TO 67 KN (15000 LBF) NEAR THE END OF THE WELD. Low-Carbon Steels to Medium-Carbon Steels. In general, low- and medium-carbon steels are joined to each other

under a wide range of conditions, and high-carbon steels are readily joined to alloy steels (Ref 7) using friction welding. High-speed tool steels are welded to alloy steel shanks for numerous machine-tool applications. Steel with carbon contents as high as 1.0%, such as 52100 steel, can be joined to lower-carbon alloys. Preweld heat treating may be required in some cases to better match the properties at the interface, and postweld heat treatment may be required in some cases to temper the interface region of the high-carbon steel grades.

Stainless Steels to Other Selected Metals. Stainless steel alloys are comparatively easy to friction weld to other

metals. For example, austenitic stainless steel to low-alloy steel (Ref 15), titanium and copper to stainless steel (Ref 16), and 1100 aluminum to stainless steel (Ref 17) are examples of transition joints that are made by FRW. Titanium can be welded to stainless steel with extreme care (Ref 18), and other incompatible dissimilar combinations may be successfully welded using interlayer techniques (Ref 19). Figure 5 shows a micrograph of the interfacial region of an inertia weld between Monel 400 and 21-6-9 stainless steel. The joint properties are excellent, with plastic flow occurring in Monel 400 before joint failure during bend testing.

FIG. 5 METALLOGRAPHIC CROSS SECTION OF THE INTERFACE OF A MONEL 400 TO 21-6-9 STAINLESS STEEL WELD PRODUCED BY INERTIA-DRIVE FRW. NOTE THE FINE GRAIN SIZE PRESENT AT THE INTERFACE.

Such transition joints can often be used as interlayers for the friction welding of incompatible materials. For example, it is difficult to weld 5083 aluminum directly to stainless steel. However, by first friction welding aluminum alloy 1100 to the stainless steel, and machining the 1100 aluminum alloy back to an interlayer thickness of about 1 mm (0.04 in.), the 5083 aluminum alloy can be joined to the stainless steel via this 1100 interlayer with high joint efficiencies (Ref 20). Figure 6(a) shows an example of an aluminum-base MMC that was friction welded to 1100 aluminum. The MMC is a 2024 aluminum alloy with 15 vol% Al2O3 particles. The interface region between these materials is shown at higher magnification in Fig. 6(b), where intermixing of both the materials is shown to occur.

FIG. 6 CROSS SECTION OF THE INTERFACE OF A DIRECT-DRIVE FRICTION WELD JOINING 1100 ALUMINUM TO A 2024 ALUMINUM ALLOY WITH 15 VOL% AL2O3 PARTICLES. (B) HIGHER MAGNIFICATION OF THE SAME WELD SHOWING THE EXCELLENT WELD FORMED AT THE INTERFACE

Problems Common to Welding of Dissimilar Materials. In general, the same problems encountered when welding

similar materials must be addressed when welding dissimilar materials. However, some problems are associated only with the welding of dissimilar materials or are greatly magnified during the welding of dissimilar materials. These factors include joint interfaces, low-melting phases, brittle phases, and different thermal expansions. Joint Interfaces. While most similar-material welds are made with little concern for surface preparation, highly

dissimilar-metal combinations are more sensitive. This happens for various reasons. In stainless steel to aluminum alloy welds, the oxide surface that forms on the aluminum picks up contaminants such as water and hydrocarbons, forming extremely tenacious surface layers (Ref 21). If this layer is not removed prior to welding, poor structural welds may occur. In stainless steel to refractory metal alloy welds, the oxide on the faying surfaces again contain contaminants such as water and hydrocarbons. The contaminants in this case are likely to alloy into the finished weldment. This alloying causes a reduction of structural integrity through the formation of low-melting or brittle phases at the weld interface. Surface-treated interfaces frequently cause problems during FRW. Steels that have been carburized or nitrided, titanium alloys that have been nitrided, and other hardfaced alloys are difficult to friction weld due to the inherently low friction coefficient and low forgeability. The repeatability of welds made on materials with hard surface layers is difficult to characterize due to several factors, including coating thickness, coating quality, and physical properties of the coating. In most instances, weldability is improved if the surface-treated area is removed from the faying surface before welding. Low-Melting Phase Formation. Some material combinations have very low melting point phases associated with

mixing of constituents at the weld interface. The formation of these phases during the welding cycle is deleterious to the finished weld properties. Examples of combinations that fall into this category include iron-base alloys to titanium alloys and aluminum alloys to magnesium alloys. Low melting point eutectics are found in both of these metallurgical systems, and great care must be exercised during parameter development to prevent the formation of liquid phases during the completion of successful welds. Other weld combinations may be affected by contaminants at the weld interface. Examples include sulfur and phosphorus in iron-base alloys and bismuth in copper alloys. These contaminants may cause problems with hot shortness in very low concentrations. It is imperative that good cleaning practices be implemented when materials may have been contaminated with these elements or with material containing these elements. Brittle Phase Formation. Many materials, when combined, are susceptible to the formation of brittle phases. In some

combinations, this occurs during the welding cycle; in others, service conditions after welding cause the problem. Two main reasons exist for brittle phase formation in friction welds: • •

SURFACE CONTAMINANTS THAT EMBRITTLE THE WELD INTERFACE (SEE THE SECTION "JOINT INTERFACES" IN THIS ARTICLE) FORMATION OF INTERMETALLIC PHASES BETWEEN NORMAL CONSTITUENTS OF THE ALLOYS BEING WELDED

Intermetallic phase formation is common when welding refractory metal alloys to stainless steel alloys and in several other systems. In the case of stainless steels to refractory metals, σ phase or similar phases may occur upon welding at the interface. Proper weld procedure development reduces the amount of brittle phases that are formed, but typically does not eliminate their formation completely. Properly developed welds have satisfactory structural properties, because only small, noncontinuous areas of the brittle phase are present at the weld interface. Figure 7(a) shows an inertia-drive welded joint between vanadium and 21-6-9 stainless steel. The interface is smooth and shows no areas of brittle phases. Electron microscopy techniques are needed to find the small areas of σ phase present at the weld interfaces.

FIG. 7 METALLOGRAPHIC CROSS SECTION OF AN INERTIA-DRIVE FRW JOINT BETWEEN VANADIUM AND A 216-9 STAINLESS STEEL. NOTE THE EXCELLENT WELD QUALITY AT THE INTERFACE. (A) WELD INTERFACE WITH NO σ-PHASE GROWTH. (B) WELD INTERFACE WITH σ-PHASE GROWTH (INDICATED BY S) AND A SOLIDSOLUTION MIXING CAUSED BY CHEMICAL DIFFUSION AFTER EXPOSURE TO A 1000 °C (1830 °F) THERMAL CYCLE

Caution must be used when designing components for use at elevated-temperature extremes. In many instances, material combinations in which no brittle phases form during welding are susceptible to brittle phase formation at the interface during high-temperature use. This is not a design issue when the welds are used for near-room-temperature applications. Figure 7(b) shows the vanadium to 21-6-9 stainless steel inertia-drive weld after a severe thermal cycle of 1000 °C (1830 °F) for 2 h with σ-phase and solid-solution growth at the weld interface. The thick layer next to the stainless steel is a solid solution of iron and vanadium. The thin layer next to the vanadium is σphase and forms a continuous brittle fracture path across the weld interface. Differential Thermal Expansion. Some material combinations are difficult to weld because of the large differences in

thermal expansion. Low-expansion materials such as refractory metals, ceramics, and low-expansion iron-nickel and ironnickel-cobalt alloys may fail or be highly stressed during cooling when welded to high-expansion material such as austenitic stainless steels and nickel-base and cobalt-base superalloys. Use of these combinations requires the designer to consider the large stresses developed within the fabricated structure if the welds are restrained when exposed to large temperature changes. Intermediate expansion materials and multiple friction welds may be required to allow for the transition from high to low thermal expansion materials.

References cited in this section

5. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 739-763 6. RECOMMENDED PRACTICES FOR FRICTION WELDING, ANSI/AWS C6. 1-89, AMERICAN WELDING SOCIETY, 1989 7. METALS HANDBOOK, 9TH ED., VOL 6, AMERICAN SOCIETY FOR METALS, 1983 8. T.N. HAZLET, PROPERTIES OF FRICTION WELDED PLAIN CARBON AND LOW ALLOY STEELS, WELD. J., VOL 41, 1962, P 49-S TO 52-S 9. S.B. DUNKERTON, TOUGHNESS PROPERTIES OF FRICTION WELDS IN STEELS, WELD. J., VOL 65, 1986, P 193-S TO 202-S 10. J.C. LIPPOLD AND B.C. ODEGARD, TECHNICAL NOTE: MICROSTRUCTURAL EVOLUTION

DURING INERTIA FRICTION WELDING OF AUSTENITIC STAINLESS STEELS, WELD. J., VOL 63, 1984, P 35-S TO 38-S 11. W.A. BAESLACK III AND K.S. HAGEY, INERTIA FRICTION WELDING OF RAPIDLY SOLIDIFIED POWDER METALLURGY ALUMINUM, WELD. J., VOL 67, 1988, P 139-S TO 149-S 12. H.H. KOO AND W.A. BAESLACK III, FRICTION WELDING OF A RAPIDLY SOLIDIFIED AL-FE-VSI ALLOY, WELD. J., VOL 71, 1992, P 147-S TO 169-S 13. K. OGAWA, H. YAMAGUCHI, T. MORIMOTO, K. TAKEMATA, H. SUDO, AND A. HIRATSUKA, SHEAR STRENGTH CHARACTERISTICS OF ALUMINUM ALLOY FRICTION WELDS, WELD. J., VOL 5, 1991, P 860-866 14. C.G. NESSLAR ET AL., FRICTION WELDING OF TITANIUM ALLOYS, WELD. J., VOL 50, 1971, P 379-S TO 385-S 15. K.G.K. MURTI AND S. SANDARESAN, THERMAL BEHAVIOR OF AUSTENITIC-FERRITIC JOINTS MADE BY FRICTION WELDING, WELD. J., VOL 64, 1985, P 327-S TO 334-S 16. R.A. BELL, J.C. LIPPOLD, AND D.R. ADOLPHSON, AN EVALUATION OF COPPER-STAINLESS STEEL FRICTION WELDS, WELD. J., VOL 63, 1984, P 325-2 TO 332-S 17. D. YASHAN, S. TSANG, W.L. JOHNS, AND M.W. DOUGHTY, INERTIA FRICTION WELDING OF 1100 ALUMINUM TO TYPE 316 STAINLESS STEEL, WELD. J., VOL 66, 1987, P 27-37 18. M. FUTAMATA AND A. FUJI, FRICTION WELDING OF TITANIUM AND SUS 304L AUSTENITIC STAINLESS STEEL, WELD. INT., VOL 4, 1990, P 768-774 19. F. SASSANI AND J.R. NEELAN, FRICTION WELDING OF INCOMPATIBLE MATERIALS, WELD. J., VOL 67, 1988, P 264-S TO 270-S 20. R. ARMSTRONG, PRIVATE COMMUNICATION AND INTERNAL REPORTS, LAWRENCE LIVERMORE NATIONAL LABORATORY, 1991 21. METALS HANDBOOK, 9TH ED., VOL 2, AMERICAN SOCIETY FOR METALS, 1979, P 204-209 Fundamentals of Friction Welding J.W. Elmer and D.D. Kautz, Lawrence Livermore National Laboratory

References

1. T.H. HAZLET, FUNDAMENTALS OF FRICTION WELDING, SOURCE BOOK ON INNOVATIVE WELDING PROCESSES, AMERICAN SOCIETY FOR METALS, 1981, P 11-36 2. F.P. BOWDEN AND D. TABOR, THE FRICTION AND LUBRICATION OF SOLIDS, PART I, OXFORD UNIVERSITY PRESS, 1954 3. J. GODDARD AND H. WILMAN, A THEORY OF FRICTION AND WEAR DURING THE ABRASION OF METALS, WEAR, VOL 5, 1962 4. V.I. VILL, FRICTION WELDING OF METALS, TRANSLATED FROM THE RUSSIAN, AMERICAN WELDING SOCIETY, 1962 5. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 739-763 6. RECOMMENDED PRACTICES FOR FRICTION WELDING, ANSI/AWS C6. 1-89, AMERICAN WELDING SOCIETY, 1989 7. METALS HANDBOOK, 9TH ED., VOL 6, AMERICAN SOCIETY FOR METALS, 1983 8. T.N. HAZLET, PROPERTIES OF FRICTION WELDED PLAIN CARBON AND LOW ALLOY STEELS, WELD. J., VOL 41, 1962, P 49-S TO 52-S 9. S.B. DUNKERTON, TOUGHNESS PROPERTIES OF FRICTION WELDS IN STEELS, WELD. J., VOL 65, 1986, P 193-S TO 202-S 10. J.C. LIPPOLD AND B.C. ODEGARD, TECHNICAL NOTE: MICROSTRUCTURAL EVOLUTION

DURING INERTIA FRICTION WELDING OF AUSTENITIC STAINLESS STEELS, WELD. J., VOL 63, 1984, P 35-S TO 38-S 11. W.A. BAESLACK III AND K.S. HAGEY, INERTIA FRICTION WELDING OF RAPIDLY SOLIDIFIED POWDER METALLURGY ALUMINUM, WELD. J., VOL 67, 1988, P 139-S TO 149-S 12. H.H. KOO AND W.A. BAESLACK III, FRICTION WELDING OF A RAPIDLY SOLIDIFIED AL-FE-VSI ALLOY, WELD. J., VOL 71, 1992, P 147-S TO 169-S 13. K. OGAWA, H. YAMAGUCHI, T. MORIMOTO, K. TAKEMATA, H. SUDO, AND A. HIRATSUKA, SHEAR STRENGTH CHARACTERISTICS OF ALUMINUM ALLOY FRICTION WELDS, WELD. J., VOL 5, 1991, P 860-866 14. C.G. NESSLAR ET AL., FRICTION WELDING OF TITANIUM ALLOYS, WELD. J., VOL 50, 1971, P 379-S TO 385-S 15. K.G.K. MURTI AND S. SANDARESAN, THERMAL BEHAVIOR OF AUSTENITIC-FERRITIC JOINTS MADE BY FRICTION WELDING, WELD. J., VOL 64, 1985, P 327-S TO 334-S 16. R.A. BELL, J.C. LIPPOLD, AND D.R. ADOLPHSON, AN EVALUATION OF COPPER-STAINLESS STEEL FRICTION WELDS, WELD. J., VOL 63, 1984, P 325-2 TO 332-S 17. D. YASHAN, S. TSANG, W.L. JOHNS, AND M.W. DOUGHTY, INERTIA FRICTION WELDING OF 1100 ALUMINUM TO TYPE 316 STAINLESS STEEL, WELD. J., VOL 66, 1987, P 27-37 18. M. FUTAMATA AND A. FUJI, FRICTION WELDING OF TITANIUM AND SUS 304L AUSTENITIC STAINLESS STEEL, WELD. INT., VOL 4, 1990, P 768-774 19. F. SASSANI AND J.R. NEELAN, FRICTION WELDING OF INCOMPATIBLE MATERIALS, WELD. J., VOL 67, 1988, P 264-S TO 270-S 20. R. ARMSTRONG, PRIVATE COMMUNICATION AND INTERNAL REPORTS, LAWRENCE LIVERMORE NATIONAL LABORATORY, 1991 21. METALS HANDBOOK, 9TH ED., VOL 2, AMERICAN SOCIETY FOR METALS, 1979, P 204-209 Fundamentals of Diffusion Bonding Murray W. Mahoney and Cliff C. Bampton, Rockwell International Science Center

Introduction DIFFUSION BONDING is only one of many solid-state joining processes wherein joining is accomplished without the need for a liquid interface (brazing) or the creation of a cast product via melting and resolidification (welding). In its most narrow definition, which is used to differentiate it from other joining processes such as deformation bonding or transient liquid phase joining, diffusion bonding (DB) is a process that produces solid-state coalescence between two materials under the following conditions: • • •

JOINING OCCURS AT A TEMPERATURE BELOW THE MELTING POINT, TM, OF THE MATERIALS TO BE JOINED (USUALLY >1/2TM ). COALESCENCE OF CONTACTING SURFACES IS PRODUCED WITH LOADS BELOW THOSE THAT WOULD CAUSE MACROSCOPIC DEFORMATION TO THE PART. A BONDING AID CAN BE USED, SUCH AS AN INTERFACE FOIL OR COATING, TO EITHER FACILITATE BONDING OR PREVENT THE CREATION OF BRITTLE PHASES BETWEEN DISSIMILAR MATERIALS, BUT THE MATERIAL SHOULD NOT PRODUCE A LOWTEMPERATURE LIQUID EUTECTIC UPON REACTION WITH THE MATERIALS TO BE JOINED.

Thus, diffusion bonding facilitates the joining of materials to produce components with no abrupt discontinuity in the microstructure and with a minimum of deformation.

Within the confines of this definition, the DB process, in practice, is limited to either press or gas pressure or bonding approaches. Illustrations and discussions of equipment and systems are presented elsewhere within this Handbook (see the Section "Solid-State Welding, Brazing, and Soldering Processes"). This article offers a qualitative summary of the theory of diffusion bonding. For those who require a more quantitative assessment, refer to the Selected References at the end of the article. It should be noted that the preferred term for this process, according to the American Welding Society, is diffusion welding. However, because diffusion bonding is used more commonly in industry, it is the term that will be used in this article. Fundamentals of Diffusion Bonding Murray W. Mahoney and Cliff C. Bampton, Rockwell International Science Center

Diffusion Bonding Process The DB process, that is, the application of pressure and temperature to an interface for a prescribed period of time, is generally considered complete when cavities fully close at the faying surfaces. Relative agreement is found for the mechanisms and sequence of events that lead to the collapse of interface voids, and the discussion below describes these metallurgical processes. Although this theoretical understanding of the DB process is universally applicable, it should be understood that parent metal strength is only approached for materials with surface conditions that do not have barriers to impede atomic bonding such as the absence of surface oxides or absorbed gases at the bonding interface. In practice, oxide-free conditions exist only for a limited number of materials. Accordingly, the properties of real surfaces limit and impede the extent of diffusion bonding. The most notable exception is titanium alloys, which, at DB temperatures greater than 850 °C (1560 °F), can readily dissolve minor amounts of adsorbed gases and thin surface oxide films and diffuse them away from the bonding surfaces, so that they will not impede the formation of the required metallic bonds across the bond interface, as shown in Fig. 1(a). An example of a successful diffusion bond in a titanium alloy is illustrated in Fig. 1(b), where the integrity of the bonded interfaces was demonstrated with subsequent superplastic expansion without interface failure.

FIG. 1 SUPERPLASTIC FORMING/DIFFUSION BONDING (SPF/DB) OF TITANIUM SHEET. (A) SEQUENCE OF OPERATIONS REQUIRED TO JOIN THREE SHEETS OF SUPERPLASTIC TITANIUM ALLOY USING SPF/DB PROCESS. (B) TYPICAL THREE-SHEET TITANIUM ALLOY COMPONENT SUPERPLASTICALLY FORMED FOLLOWING DIFFUSION BONDING.

Similarly, the joining of silver at 200 °C (390 °F) requires no deformation to break up and disperse oxides, because silver oxide dissociates completely at 190 °C (375 °F). Above this temperature, silver dissolves its oxide and also scavenges many surface contaminants. Other examples of metals that have a high solubility for interstitial contaminants include tantalum, tungsten, copper, iron, zirconium, and niobium. Accordingly, this class of alloy is easiest to diffusion bond. A second class of material, that is, metals and alloys that exhibit very low solubility for interstitials (such as aluminum-, iron-, nickel-, and cobalt-base alloys) are not readily diffusion bondable. Special consideration must be given to remove surface barriers to atomic diffusion prior to joining and subsequently prevent their reformation during the joining process. This is not an easy processing matter. Accordingly, the potential for high-strength bond interfaces for alloys with low interstitial solubility should be considered on an individual alloy basis. Fundamentals of Diffusion Bonding Murray W. Mahoney and Cliff C. Bampton, Rockwell International Science Center

Bonding Surfaces Containing Oxides Diffusion bonding can be achieved for materials with adherent surface oxides, but the resultant interface strengths of these materials are considerably less than that measured for the parent material. Aluminum alloys are prime examples of this class of material. Research since 1960 has demonstrated only limited diffusion bond properties. Although interface

strength can be increased for oxide-bearing materials, it requires considerable surface extension of the faying interfaces to create localized plastic flow of the metal and concurrent oxide breakage. This introduces an increased number of locations for metal-to-metal contact via plastic flow around or microextrusion through the broken oxide. In general, the oxide is not removed, but is simply dispersed over a greater surface area in an enclosed environment, in which oxidation cannot recur. Thus, even with significant surface deformation, only a fraction of the interface area contributes to the strength of the bond. The proportion of oxide-free metallic area revealed is dependent on the relative hardness of the metal and its oxide film, as well as on the mechanical properties of the oxide. This type of bonding, although often considered as diffusion bonding, is better described as deformation bonding and does not fit within the strict definition of the low deformation associated with diffusion bonding. Factors that affect the relative difficulty of diffusion bonding oxide-bearing surfaces include: • • •

•

SURFACE ROUGHNESS PRIOR TO WELDING. A ROUGHER SURFACE WILL RESULT IN GREATER SHEAR DEFORMATION. MECHANICAL PROPERTIES OF THE OXIDE. THE MORE BRITTLE THE OXIDE, THE GREATER THE DISPERSION FOR A GIVEN LEVEL OF DEFORMATION. RELATIVE HARDNESS OF THE METAL AND ITS OXIDE FILM. BECAUSE PLASTIC FLOW CONTROLS THE AMOUNT OF BONDING AREA, LARGE DIFFERENCES IN THEIR HARDNESS SHOULD FACILITATE BONDING. PRESTRAINING OR WORK HARDENING OF THE MATERIAL. INITIATION OF BONDING WILL OCCUR AT LOWER DEFORMATIONS FOR PRESTRAINED OR WORK-HARDENED MATERIALS, AND THE DEGREE OF SURFACE EXTENSION IN THE CENTRAL REGION OF THE INTERFACE IS CONSIDERABLY GREATER FOR COLD-WORKED MATERIAL. THUS, ANNEALED MATERIAL REQUIRES A LARGER TOTAL DEFORMATION BEFORE BONDING WILL INITIATE.

It is clear that with the appropriate information, sufficient experiments can be performed to determine the diffusion bondability of most materials. Parent metal strength will not always be attained using the DB approach, particularly for materials with adherent oxides, but interface strength can be maximized if the fundamentals of the process are understood. Fundamentals of Diffusion Bonding Murray W. Mahoney and Cliff C. Bampton, Rockwell International Science Center

Mechanism of Diffusion Bonding In diffusion bonding, the nature of the joining process is essentially the coalescence of two atomically clean solid surfaces. Complete coalescence comes about through a three-stage metallurgical sequence of events. Each stage, as shown in Fig. 2, is associated with a particular metallurgical mechanism that makes the dominant contribution to the bonding process. Consequently, the stages are not discretely defined, but begin and end gradually, because the metallurgical mechanisms overlap in time. During the first stage, the contact area grows to a large fraction of the joint area by localized deformation of the contacting surface asperities. Factors such as surface roughness, yield strength, work hardening, temperature, and pressure are of primary importance during this stage of bonding. At the completion of this stage, the interface boundary is no longer a planar interface, but consists of voids separated by areas of intimate contact. In these areas of contact, the joint becomes equivalent to a grain boundary between the grains on each surface. The first stage is usually of short duration for the common case of relatively high-pressure diffusion bonding.

FIG. 2 SEQUENCE OF METALLURGICAL STAGES IN DIFFUSION BONDING PROCESS. (A) INITIAL CONTACT: LIMITED TO A FEW ASPERITIES (ROOM TEMPERATURE). (B) FIRST STAGE: DEFORMATION OF SURFACE ASPERITIES BY PLASTIC FLOW AND CREEP. (C) SECOND STAGE: GRAIN BOUNDARY DIFFUSION OF ATOMS TO THE VOIDS AND GRAIN BOUNDARY MIGRATION. (D) THIRD STAGE: VOLUME DIFFUSION OF ATOMS TO THE VOIDS

During the second stage of joint formation, two changes occur simultaneously. All of the voids in the joints shrink, and most are eliminated. In addition, the interfacial grain boundary migrates out of the plane of the joint to a lower-energy equilibrium. Creep and diffusion mechanisms are important during the second stage of bonding and for most, if not all, practical applications, bonding would be considered essentially complete following this stage. As the boundary moves, any remaining voids are engulfed within grains where they are no longer in contact with a grain boundary. During this third stage of bonding, the voids are very small and very likely have no impact on interface strength. Again, diffusional processes cause the shrinkage and elimination of voids, but the only possible diffusion path is now through the volume of the grains themselves. Stage I: Microasperity Deformation

The nature of the starting surface is of considerable importance, because of the small macroscopic deformation allowed during diffusion bonding. A real surface is never perfectly clean or perfectly smooth, and the area of metal-to-metal contact between faying surfaces is a very small fraction of the area of joint contact. Contact is limited to a relatively few microasperities. At room temperature and under load, these asperities deform as long as the surface area of contact is such that the yield strength of the material is exceeded. The extent of this deformation is limited at room temperature, and is even more limited for work-hardenable materials. As temperature increases to the diffusion bonding temperature, the flow stress of the material decreases and additional asperity deformation occurs through plastic flow. Again, flow occurs until the area of contact increases to an extent that the yield strength of the material is exceeded. If the temperature is above the recrystalization temperature of the material, then work hardening is no longer a consideration. With time at temperature, creep mechanisms now control the rate of asperity deformation, and the area of contact or bond continues to grow. As the area of contract grows, the stress acting on the surface asperities decreases. Consequently, creep deformation progressively slows and diminishes in significance. The contributions of temperature and pressure to both plastic and creep deformation during this initial stage of diffusion bonding are synergistic, that is, at higher temperatures, less pressure is required and vice versa. However, for any combination of temperature and pressure, bulk deformation to the part is limited to a small percentage ( 5) than for the tensile specimens (εf ~0.001). The shorter rupture times for tensile joints are a direct result of the hydrostatic tensile stresses superimposed on the effective stresses caused by the mechanical constraint imposed by the base metal.

FIG. 9 COMPARISON BETWEEN THE RUPTURE TIME VERSUS EFFECTIVE (VON MISES) STRESS BEHAVIOR IN SHEAR, AT TORSIONAL STRESSES LESS THAN THE MAXIMUM OF HHC-DEPOSITED SILVER INTERLAYERS, COMPARED WITH THAT OF TENSION, CALCULATED USING FEM ANALYSES OF PM SPUTTER-DEPOSITED SILVER INTERLAYERS

The change in steady-state strain-rate (or time-to-rupture, according to the Monkman-Grant relationship) with applied stress is predictable according to the steady-state stress exponent, n,  ∂ ln ε  n=   ∂ ln σ 

(EQ 2)

where n is about 30 for silver at ambient temperature (Ref 38). Naturally, the steady-state strain rate of the torsional specimens shown in Fig. 9 is less than the constant applied strain rate of the torsional specimens shown in Fig. 8. If the steady-state rate if not achievable in the interlayer metal, then time-dependent failure in shear is not expected to occur. It is important to realize the consequence of the difference in plastic strains-to-failure between the constant shear-stress case and the constant tensile-stress case. For example, residual shear or tensile stresses may develop during fabrication of interlayer brazes or solid-state welds. This is particularly true in cases where dissimilar base materials are joined, but have substantially different coefficients of thermal expansion. For thin interlayers subjected to residual shear stresses, the dimensions of the base metal and interlayer may be such that the stress is relaxed at strains less than the shear strain-tofailure. However, because the observed tensile strains-to-failure are small, creep rupture by residual tensile stresses is possible, provided a significant stress level exists. If the interlayer metal is resistant to tim-dependent failure, then only the indirect implications of residual stress are important (for example, environmentally induced failure).

References cited in this section

17. R.S. ROSEN, "TIME-DEPENDENT FAILURE OF SILVER INTERLAYER WELDS," PH.D.THESIS, UNIVERSITY OF CALIFORNIA, DAVIS, 1990 26. J.W. ELMER, M.E. KASSNER, AND R.S. ROSEN, THE BEHAVIOR OF SILVER-AIDED DIFFUSIONWELDED JOINTS UNDER TENSILE AND TORSIONAL LOADS, WELD. J., VOL 67 (NO. 7), 1988, P 157S-162S 33. M.E. KASSNER, R.S. ROSEN, AND G.A. HENSHALL, DELAYED MECHANICAL FAILURE OF SILVER-INTERLAYER DIFFUSION BONDS, METALL. TRANS. A, VOL 21, 1990, P 3085-3100 38. M.E. KASSNER, THE RATE DEPENDENCE AND MICROSTRUCTURE OF HIGH-PURITY SILVER DEFORMED TO LARGE STRAINS BETWEEN 0.16 AND 0.30 TM, METALL. TRANS. A, VOL 20, 1989, P 2001-2010 Mechanical Properties of Soft-Interlayer Solid-State Welds R.S. Rosen, Lawrence Livermore National Laboratory; M.E. Kassner, Oregon State University

Multiaxial Loading As discussed earlier, in tension there is significant mechanical constraint by the base material, which tends to reduce the effective stress. In torsion, where the shear stress is parallel to the plane of the interlayer, there is not constraint, other than, perhaps, the relatively minor constraint associated with plastic strain incompatibilities between the interlayer and the base metal. It was also mentioned earlier that the equivalent uniaxial strain-to-failure for torsional deformation of softinterlayer joints is roughly four orders of magnitude higher than the values for uniaxial tension, and the (effective) shear stress-to-failure is only about one-half the UTS. This suggests that the interlayers will behave anisotropically to imposed stresses. The mechanical properties of the intermediate cases of biaxial deformation are not readily interpolated between these limits. The authors have preliminary experimental evidence, for example, that low-strain ductile failures in thin interlayers may occur at particularly low stress levels and strains for cases where tensile loads are accompanied by inplane shear stresses (Ref 63). This implies that some multiaxial residual stress states may render solid-state-welded interlayer joints substantially more vulnerable to failure than simple tensile-stress states, even though the "macroscopic" effective stress may be comparable, when calculated on the basis of the applied stress state in the base material away from the interlayer.

Reference cited in this section

63. M.E. KASSNER, OREGON STATE UNIVERSITY, UNPUBLISHED RESEARCH, 1989

Mechanical Properties of Soft-Interlayer Solid-State Welds R.S. Rosen, Lawrence Livermore National Laboratory; M.E. Kassner, Oregon State University

Environmentally Induced Failure of Interlayers The interlayer/base-metal interfaces of some soft-interlayer welds appear particularly vulnerable to failure when exposed to external (corrosive) environments. For example, it is known that brazed silver interlayers between stainless steel are subject to galvanic corrosion in aqueous NaCl media (Ref 64, 65). When dissimilar base materials are joined, the potential for stress-corrosion cracking (SCC) exists if the joints are exposed simultaneously to some external (corrosive) environments and either shear or tensile stresses. If an interlayer is subjected to stress, then a determination of the vulnerability to SCC is appropriate. It was mentioned earlier that elevated-temperature joining using either solid-state or brazing processes may leave the joint under substantial residual stress, particularly in the case of dissimilar metals for which the coefficients of thermal expansion are significantly different. The authors have reported (Ref 35) that the uranium to type 304 stainless steel solid-state welds, fabricated using HHC-deposited silver interlayers (Ref 25, 26), are very susceptible to SCC. These silver solid-state welds were joined at 873 K, and the difference in thermal expansion coefficients leaves the joints of cylindrical specimens under significant residual shear stresses after cooling to ambient temperature. It was found that these residual stresses led to SCC at the uranium/silver interface in air saturated with water vapor (100% relative humidity) (Ref 35). This cracking resulted in reduced tensile strengths and rupture times, when compared with those specimens tested in laboratory air ( 40% relative humidity). Specimens subjected to the same 100% relative humidity in air in an unstressed state exhibited no loss in strength, even after lengthy exposure. This, or course, emphasizes the importance of considering the residual stresses as not only rendering the interlayer joints vulnerable to "pure mechanical" rupture, but also to environmentally induced failures such as SCC, hydrogen embrittlement, and metal-induced embrittlement, if exposed to critical environments.

References cited in this section

25. R.S. ROSEN, D.R. WALMSLEY, AND Z.A. MUNIR, THE PROPERTIES OF SILVER-AIDED DIFFUSION WELDS BETWEEN URANIUM AND STAINLESS STEEL, WELD. J., VOL 65 (NO. 4), 1986, P 83S-92S 26. J.W. ELMER, M.E. KASSNER, AND R.S. ROSEN, THE BEHAVIOR OF SILVER-AIDED DIFFUSIONWELDED JOINTS UNDER TENSILE AND TORSIONAL LOADS, WELD. J., VOL 67 (NO. 7), 1988, P 157S-162S 35. R.S. ROSEN, S. BEITSCHER, AND M.E. KASSNER, STRESS CORROSION CRACKING OF URANIUM-SILVER INTERFACES IN SILVER-AIDED DIFFUSION WELDS, INT. CONF. ENVIRONMENT-INDUCED CRACKING OF METALS, VOL 10, NACE, 1989, P 429-433 64. A.T. KUHN AND R.M. TRIMMER, REVIEW OF THE AQUEOUS CORROSION OF STAINLESS STEEL-SILVER BRAZED JOINTS, BR. CORROS. J., VOL 17 (NO. 1), 1982, P 4-8 65. T. TAKEMOTO AND I. OKAMOTO, EFFECT OF COMPOSITION ON THE CORROSION BEHAVIOR OF STAINLESS STEELS BRAZED WITH SILVER-BASE FILLER METALS, WELD. J., VOL 64, 1984, P 300S-307S Mechanical Properties of Soft-Interlayer Solid-State Welds R.S. Rosen, Lawrence Livermore National Laboratory; M.E. Kassner, Oregon State University

References

1. J.T. NIEMANN, R.P. SOPHER, AND P.J. RIEPPEL, DIFFUSION BONDING BELOW 1000 °F, WELD. J., VOL 37, 1958, P 337S-342S 2. I.M. BARTA, LOW TEMPERATURE DIFFUSION BONDING OF ALUMINUM ALLOYS, WELD. J., VOL 43, 1964, P 241S-247S

3. D. HAUSER, P.A. KAMMER, AND J.H. DEDRICK, SOLID-STATE WELDING OF ALUMINUM, WELD. J., VOL 46, 1967, P 11S-22S 4. P.A. KAMMER, R.E. MONROE, AND D.C. MARTIN, FURTHER STUDIES OF DIFFUSION BONDING BELOW 1000 °F, WELD. J., VOL 48, 1969, P 116S-124S 5. Y. IINO AND N. TAGAUCHI, INTERDIFFUSING METALS LAYER TECHNIQUE OF CERAMICMETAL BONDING, J. MATER. SCI. LETT., VOL 7 (NO. 9), 1988, P 981-982 6. A. URENA, J.M.G.D. SALAZAR, AND J. QUINONES, DIFFUSION BONDING OF ALUMINA TO STEEL USING SOFT COPPER INTERLAYER, J. MATER. SCI., VOL 27, 1992, P 599-606 7. T. YAMADA K, YOKOI, AND A. KOHNO, EFFECT OF RESIDUAL STRESS ON THE STRENGTH OF ALUMINA-STEEL JOINT WITH AL-SI INTERLAYER, J. MATER. SCI., VOL 25, 1990, P 21882192 8. M.G. NICHOLAS AND R.M. CRISPIN, DIFFUSION BONDING STAINLESS STEEL TO ALUMINA USING ALUMINUM INTERLAYERS, J. MATER. SCI., VOL 17, 1982, P 3347-3360 9. C.L. CLINE, AN ANALYTICAL AND EXPERIMENTAL STUDY OF DIFFUSION BONDING, WELD. J., VOL 45, 1966, P 481S-489S 10. A.T. D'ANNESSA, THE SOLID-STATE BONDING OF REFRACTORY METALS, WELD. J., VOL 43, 1964, P 232S-240S 11. M.G. NICHOLAS AND R.M. CRISPIN, DIFFUSION BONDING CERAMICS WITH DUCTILE METAL INTERLAYERS, PROC. INT. DIFFUSION BONDING (CRANFIELD, U.K.), 1987, P 173-182 12. B. DERBY, ZIRCONIA/METAL DIFFUSION BONDS, PROC. INT. CONF. DIFFUSION BONDING (CRANFIELD, U.K.), 1987, P 195-201 13. R.A. MORLEY AND J. CARUSO, THE DIFFUSION WELDING OF 390 ALUMINUM ALLOY HYDRAULIC VALVE BODIES, WELD. J., VOL 59 (NO. 8), 1980, P 29-34 14. C.H. CRANE, D.T. LOVELL, W.A. BAGINSKY, AND M.G. OLSEN, DIFFUSION WELDING OF DISSIMILAR METALS, WELD. J., VOL 46, 1967, P 23S-31S 15. M. O'BRIEN, C.R. RICE, AND D.L. OLSON, HIGH STRENGTH DIFFUSION WELDING OF SILVER COATED BASE METALS, WELD. J., VOL 55 (NO. 1), 1976, P 25-27 16. J.W. DINI, W.K. KELLEY, W.C. COWDEN, AND E.M. LOPEZ, USE OF ELECTRODEPOSITED SILVER AS AN AID IN DIFFUSION WELDING, WELD. J., VOL 63 (NO. 1), 1983, P 26S-34S 17. R.S. ROSEN, "TIME-DEPENDENT FAILURE OF SILVER INTERLAYER WELDS," PH.D.THESIS, UNIVERSITY OF CALIFORNIA, DAVIS, 1990 18. J.L. KNOWLES AND T.H. HAZLETT, HIGH-STRENGTH LOW-TEMPERATURE BONDING OF BERYLLIUM AND OTHER METALS, WELD. J., VOL 49 (NO. 7), 1970, P 301S-310S 19. P.S. MCLEOD AND G. MAH, THE EFFECT OF SUBSTRATE BIAS VOLTAGE ON THE BONDING OF EVAPORATED SILVER COATINGS, J. VAC. SCI. TECHNOL., VOL 11 (NO. 1), 1974, P 119-121 20. G. MAH, P.S. MCLEOD, AND D.G. WILLIAMS, CHARACTERIZATION OF SILVER COATINGS DEPOSITED FROM A HOLLOW CATHODE SOURCE, J. VAC. SCI. TECHNOL., VOL 11 (NO. 4), 1974, P 663-665 21. D.G. WILLIAMS, VACUUM COATING WITH A HOLLOW CATHODE SOURCE, J. VAC. SCI. TECHNOL., VOL 11 (NO. 1), 1974, P 374-377 22. E.R. NAIMON, D. VIGIL, J.P. VILLEGAS, AND L. WILLIAMS, ADHESION STUDY OF SILVER FILMS DEPOSITED FROM A HOY HOLLOW-CATHODE SOURCE, J. VAC. SCI TECHNOL., VOL 13 (NO. 6), 1976, P 1131-1135 23. E.R. NAIMON, J.H. DOYLE, C.R. RICE, D. VIGIL, AND D.R. WALMSLEY, DIFFUSION WELDING OF ALUMINUM TO STAINLESS STEEL, WELD. J., VOL 60 (NO. 1), 1981, P 17-20 24. D.T. LARSON AND H.L. DRAPER, CHARACTERIZATION OF THE BE-AG INTERFACIAL REGION OF SILVER FILMS DEPOSITED ONTO BERYLLIUM USING A HOT HOLLOW CATHODE DISCHARGE, THIN SOLID FILMS, VOL 107, 1983, P 327-334 25. R.S. ROSEN, D.R. WALMSLEY, AND Z.A. MUNIR, THE PROPERTIES OF SILVER-AIDED

DIFFUSION WELDS BETWEEN URANIUM AND STAINLESS STEEL, WELD. J., VOL 65 (NO. 4), 1986, P 83S-92S 26. J.W. ELMER, M.E. KASSNER, AND R.S. ROSEN, THE BEHAVIOR OF SILVER-AIDED DIFFUSIONWELDED JOINTS UNDER TENSILE AND TORSIONAL LOADS, WELD. J., VOL 67 (NO. 7), 1988, P 157S-162S 27. J. HARVEY, P.G. PATRIDGE, AND A.M. LURSHEY, FACTORS AFFECTING THE SHEAR STRENGTH OF SOLID STATE DIFFUSION BONDS BETWEEN SILVER-COATED CLAD AL-ZNMG ALLOY (ALUMINUM ALLOY 70100, MATER. SCI. ENG., VOL 79, 1986, P 191-199 28. P.G. PARTRIDGE AND D.V. DUNFORD, ON THE TESTING OF DIFFUSION-BONDED OVERLAP JOINTS BETWEEN CLAD AL-ZN-MG ALLOY (7010) SHEET, J. MATER. SCI. VOL 22, 1987 P 15971608 29. D.V. DUNFORD AND P.G. PARTRIDGE, THE PEEL STRENGTHS OF DIFFUSION BONDED. JOINTS BETWEEN CLAD AL-ALLOY SHEETS, J. MATER. SCI., VOL 22, 1987, P 1790-1798 30. R.S. ROSEN AND M.E. KASSNER, DIFFUSION WELDING OF SILVER INTERLAYERS COATED ONTO BASE METALS BY PLANAR-MAGNETRON SPUTTERING, J. VAC. SCI. TECHNOL. A, VOL 8 (NO. 1), 1900, P 19-29 31. M.E. KASSNER, R.S. ROSEN, G.A. HENSHALL, AND W.E. KING, DELAYED FAILURE OF SILVER AIDED DIFFUSION WELDS BETWEEN STEEL, PROC. 2ND INT. CONF. BRAZING, HIGH TEMPERATURE BRAZING, AND DIFFUSION WELDING, GERMAN WELDING SOCIETY, 1989, P 4752 32. M.E. KASSNER, R.S. ROSEN, G.A. HENSHALL, AND K.D. CHALLENGER, TIME-DEPENDENT FAILURE OF SILVER-INTERLAYER DIFFUSION WELDS BETWEEN ELASTICALLYDEFORMING BASE METALS, SCR. METALL. MATER., VOL 24, 1990, P 587-592 33. M.E. KASSNER, R.S. ROSEN, AND G.A. HENSHALL, DELAYED MECHANICAL FAILURE OF SILVER-INTERLAYER DIFFUSION BONDS, METALL. TRANS. A, VOL 21, 1990, P 3085-3100 34. P.G. PARTRIDGE AND J. HARVEY, "PROCESS FOR THE DIFFUSION BONDING OF ALUMINUM MATERIALS," U.K. PATENT APPLICATIONS GB 2117691 A, 1983 35. R.S. ROSEN, S. BEITSCHER, AND M.E. KASSNER, STRESS CORROSION CRACKING OF URANIUM-SILVER INTERFACES IN SILVER-AIDED DIFFUSION WELDS, INT. CONF. ENVIRONMENT-INDUCED CRACKING OF METALS, VOL 10, NACE, 1989, P 429-433 36. E. OROWAN, FRACTURE AND STRENGTH OF SOLIDS, REP. PROG. PHYS., VOL 12, 1948, P 185231 37. E. OROWAN, FUNDAMENTALS OF BRITTLE BEHAVIOR OF METALS, FATIGUE AND FRACTURE OF METALS, JOHN WILEY & SONS, 1952 38. M.E. KASSNER, THE RATE DEPENDENCE AND MICROSTRUCTURE OF HIGH-PURITY SILVER DEFORMED TO LARGE STRAINS BETWEEN 0.16 AND 0.30 TM, METALL. TRANS. A, VOL 20, 1989, P 2001-2010 39. N. BREDZ, INVESTIGATION OF FACTORS DETERMINING THE TENSILE STRENGTH OF BRAZED JOINTS, WELD. J., VOL 33, 1954, P 545S-563S 40. W.G. MOFFATT AND J. WULFF, TENSILE DEFORMATION AND FRACTURE OF BRAZED JOINTS, WELD. J., VOL 42, 1963, P 115S-125S 41. H.J. SAXTON, A.J. WEST, AND C.R. BARRETT, DEFORMATION AND FAILURE OF BRAZED JOINTS--MACROSCOPIC CONSIDERATIONS, METALL. TRANS., VOL 2, 1971, P 999-1007 42. A.J. WEST, H.J. SAXTON, A.S. TETELMAN, AND C.R. BARRETT, DEFORMATION AND FAILURE OF THIN BRAZED JOINTS--MICROSCOPIC CONSIDERATIONS, METALL. TRANS., VOL 2, 1971, P 1009-1017 43. E.P. LAUTENSCHLAGER, B.C. MARKER, B.K. MOORE, AND R. WILDES, STRENGTH MECHANISMS OF DENTAL SOLDER JOINTS, J. DENT. RES., VOL. 53, 1974, P 1361-1367 44. R.A. MUSIN, V.A. ANTSIVEROV, Y.A. BELIKOV, Y.V. LYAMIN, AND A.N. SOLOKOV, THE EFFECTS ON THE STRENGTH OF THE JOINT OF THE THICKNESS OF THE SOFT INTERLAYER

IN DIFFUSION WELDING, AUTO. WELD., VOL 32, 1979, P 38-40 45. R.M. TRIMMER AND A.T. KUHN, THE STRENGTH OF SILVER-BRAZED STEEL JOINTS--A REVIEW, BRAZING SOLDERING, VOL 2, 1982 P 6-13 46. N. BREDZ AND H. SCHWARTZBART, TRIAXIAL TENSION TESTING AND THE BRITTLE FRACTURE STRENGTH OF METALS, WELD. J., VOL 53 (NO. 12), 1956, P 610S-615S 47. W.G. MOFFATT AND J. WULFF, STRENGTH OF SILVER BRAZED JOINTS IN MILD STEEL, TRANS. AIME, VOL 209 (NO. 4), 1957, P 442-445 48. R.Z. SHRON AND O.A. BAKSHI, THE PROBLEMS OF GAUGING THE STRENGTHS OF WELDED JOINTS IN WHICH THERE IS A SOFT INTERLAYER, WELD. PROD., VOL 9, 1962, P 19-23 49. O.A. BAKSHI AND A.A. SHATOV, THE STRESSED STATE IN WELDED JOINTS WITH HARD AND SOFT INTERLAYERS, WELD PROD., VOL 13, 1966, P 13-19 50. V.S. GOLOVCHENKO AND B.A. GOLOBOV, THE USE OF SILVER AS AN INTERMEDIATE LAYER FOR JOINING TITANIUM TO OTHER METALS, WELD. PROD., VOL 18, 1971, P 55-57 51. G.K. KHARCHENKO AND A.I. IGNATENKO, THE STRENGTH OF JOINTS WITH A THIN SOFT INTERLAYER, AUTO. WELD., VOL 21, 1968, P 33-35 52. W.M. LEHER AND H. SCHWARTZBART, STATIC AND FATIGUE STRENGTHS OF METALS SUBJECTED TO TRIAXIAL STRESSES, VOL 60, 1960, P 610-626 53. E.A. ALMOND, D.K. BROWN, G.J. DAVIES, AND A.M. COTTENDED, THEORETICAL AND EXPERIMENTAL INTERLAYED BUTT JOINTS TESTED IN TENSION, INT. J. MECH. SCI., VOL 25 (NO. 3), 1983, P 175-189 54. G.A. HENSHALL, R.S. ROSEN, M.E. KASSNER, AND R.G. WHIRLEY, FINITE ELEMENT ANALYSIS OF INTERLAYER WELDS LOADED IN TENSION, WELD. J., VOL 69 (NO. 9) 1990, P 337S-345S 55. H.J. SAXTON, A.J. WEST, AND C.R. BARRETT, THE EFFECT OF COOLING RATE ON THE STRENGTH OF BRAZED JOINTS, METALL. TRANS., VOL 2, 1971, P 1019-1028 56. M.C. TOLLE AND M.E. KASSNER, TENSILE PROPERTIES OF THIN AU-NI BRAZES BETWEEN STRONG BASE METALS, SCR. METALL. MATER., VOL 26, 1992, P 1281-1284 57. M.E. KASSNER, M.C. TOLLE. R.S. ROSEN, G.A. HENSHALL, AND J.W. ELMER, RECENT ADVANCES IN UNDERSTANDING THE MECHANICAL BEHAVIOR OF CONSTRAINED THIN METALS IN BRAZES AND SOLID-STATE BONDS, THE METAL SCIENCE OF JOINING, M.J. CIESLAK, ED., TMS, 1992, P 223-232 58. R.W. LOGAN, R.G. CASTRO, AND A.K. MUKHERJEE, MECHANICAL PROPERTIES OF SILVER AT LOW TEMPERATURES, SCR. METALL., VOL 17, 1983, P 63-66 59. T.J. MOORE AND K.H. HOLKO, SOLID-STATE WELDING OF TD-NICKEL BAR, WELD. J., VOL 49 (NO. 9), 1970, P 395S-409S 60. R.S. ROSEN, "AN INVESTIGATION OF THE PROPERTIES OF A SILVER-AIDED SOLID-STATE BOND BETWEEN URANIUM AND STAINLESS STEEL," REPORT UCRL-53458, LAWRENCE LIVERMORE NATIONAL LABORATORY, 1983 61. E.R. NAIMON, R.G. KURZ, D. VIGIL, AND L. WILLIAMS, "SILVER FILMS FOR SOLID STATE BONDING,"REPORT RFP-3125, ROCKWELL INTERNATIONAL, 1981 62. Z. NISENHOLTZ, J. MIRONI, AND N. NIR, DIFFUSION BONDING OF STAINLESS STEEL 304L TO TI-6AL-4V ALLOY, PROC. 3RD INT. CONF. ISOSTATIC PRESSING (LONDON, U.K.), 1986 63. M.E. KASSNER, OREGON STATE UNIVERSITY, UNPUBLISHED RESEARCH, 1989 64. A.T. KUHN AND R.M. TRIMMER, REVIEW OF THE AQUEOUS CORROSION OF STAINLESS STEEL-SILVER BRAZED JOINTS, BR. CORROS. J., VOL 17 (NO. 1), 1982, P 4-8 65. T. TAKEMOTO AND I. OKAMOTO, EFFECT OF COMPOSITION ON THE CORROSION BEHAVIOR OF STAINLESS STEELS BRAZED WITH SILVER-BASE FILLER METALS, WELD. J., VOL 64, 1984, P 300S-307S

Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Introduction SHIELDED METAL ARC WELDING (SMAW), commonly called stick, or covered electrode, welding, is a manual welding process whereby an arc is generated between a flux-covered consumable electrode and the workpiece. The process uses the decomposition of the flux covering to generate a shielding gas and to provide fluxing elements to protect the molten weld-metal droplets and the weld pool. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

The SMAW Process The important features of the SMAW process are shown in Fig. 1. The arc is initiated by momentarily touching or "scratching" the electrode on the base metal. The resulting arc melts both the base metal and the tip of the welding electrode. The molten electrode metal/flux is transferred across the arc (by arc forces) to the base-metal pool, where it becomes the weld deposit covered by the protective, less-dense slag from the electrode covering.

FIG. 1 SMAW PROCESS

Advantages and Limitations. The SMAW process is the most widely used welding process. It is the simplest, in terms

of equipment requirements, but it is, perhaps, the most difficult in terms of welder training and skill-level requirements. Although welder skill level is a concern, most welders entering the field start as "stick welders" and develop the necessary skills through training and experience. The equipment investment is relatively small, and welding electrodes (except the very reactive metals, such as titanium, magnesium, and others) are available for virtually all manufacturing, construction, or maintenance applications. Shielded metal arc welding has the greatest flexibility of all the welding processes, because it can be used in all positions (flat, vertical, horizontal, and overhead), with virtually all base-metal thicknesses (1.6 mm, or

1 in., and greater), and in areas of limited accessibility, which is a very important capability. 16

Because the SMAW process is basically a manual process, the skill level of the welder is of paramount importance in obtaining an acceptable weld. The welder duty cycle is generally low, because of the built-in work break, which occurs after each electrode is consumed and requires replacement. In addition to replacing the electrode when the arc is stopped (broken), the welder may "chip" or remove slag and clean it away from the starting and welding area with a wire brush to allow the proper deposition of the subsequent weld. This electrode replacement and cleaning operation occurs many times during the work day (about every two minutes, or the time it generally takes to consume an electrode). This stopping,

chipping, wire brushing, and electrode replacement prevents the welder from attaining an operator factor, or duty cycle, that is much greater than 25%. Weld Quality. The quality of the weld depends on the design and accessibility of the joint, as well as on the electrode, the technique, and the skill of the welder. If joint details vary greatly from established design details, then a lower-quality weld can result. Other factors that also reduce quality are improper interbead cleaning, poor location of individual weld beads within the joint, and various problems with individual electrodes, including partially missing flux and core wires that are not centered within the flux covering. Overall, welds of excellent quality can be obtained with the SMAW process, as demonstrated by its use in joining submarine pressure hull sections and high-pressure oil/gas pipe lines. The base-metal thicknesses. that can be welded using the SMAW process generally range from 1.6 mm (

1 in.) to an 16

unlimited thickness. The thinner materials require a skilled welder, tight fitup, and the proper small-diameter welding electrode. Welding position also is important when determining the minimum plate thicknesses that can be welded. Flatposition butt welds and horizontal fillet welds are generally considered the easiest to weld. Out-of-position welding (vertical, overhead) requires greater skill. Welding Circuit. The circuit diagram for the SMAW process is shown in Fig. 2. The equipment consists of a power

source, electrode holder, and welding cables that connect the power source to the electrode holder and the workpiece. Alternating current (ac), or direct current, electrode negative (DCEN), or direct current, electrode positive (DCEP) can be used, depending on the electrode coating characteristics. The DCEN source is also called dc straight polarity, whereas the DCEP source is also called dc reverse polarity.

FIG. 2 SHIELDED METAL ARC CIRCUIT DIAGRAM

Equipment. The welding machine, or power source, is the crux of the SMAW process. Its primary purpose is to provide electrical power of the proper current and voltage to maintain a controllable and stable welding arc. Its output characteristics must be of the constant current (CC) type. SMAW electrodes operate within the range from 25 to 500 A. The electrode producer should suggest a narrow optimum range for each size and type of electrode. Operating arc voltage varies between 15 and 35 V.

The electrode holder, which is held by the welder, firmly grips the electrode and transmits the welding current to it. Electrode holders are available in several designs, such as the pincher type and the collet, or twist, type, shown in Fig. 3. Each style has its proponents and the selection is usually a personal preference. Electrode holders are designated by their current capacity. Selection factors, such as the current rating, duty cycle, maximum electrode size, and cable size, are shown in Table 1. The most lightweight holder that will accommodate the required electrode size is usually desired.

TABLE 1 SIZE AND CAPACITY OF ELECTRODE HOLDERS ELECTRODE HOLDER

RATING

MAXIMUM

MAXIMUM

ELECTRODE SIZE

CABLE SIZE

50

mm 3.2

1

200

50

4.0

MEDIUM

300

60

5.5

LARGE

400

60

6.4

EXTRA LARGE

500

75

7.9

600

75

9.5

SMALL

MAXIMUM CURRENT, A

DUTY CYCLE, %

100

in. 1 8 5 32 7 32 1 4 5 16 3 8

1/0 2/0 3/0 4/0 4/0

FIG. 3 SMAW ELECTRODE HOLDERS

All electrode holders should be fully insulated. Because they are used in proximity to the arc and are exposed to high heat, they will deteriorate rapidly. It is extremely important to maintain electrode holders to ensure that they retain their current-carrying efficiency, their insulating qualities, and their electrode gripping action. Manufacturers supply spare parts so that the holders can be rebuilt and maintained for safe and efficient operation. Certain pieces of auxiliary equipment can be used with the SMAW process, such as low-voltage control circuits, which enable the relatively high open-circuit voltage to be cut off until the electrode touches the workpiece. Other items include remote-control switches for the contactors, remote-control current adjusting devices, and engine idling controllers for engine-driven power sources. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Applications Most manufacturing operations that require welding will strive to utilize the mechanized processes that offer greater productivity, higher quality, and, therefore, more cost-effective production. For these reasons, the SMAW process has been replaced where possible. However, the simplicity and ability of the SMAW process to achieve welds in areas of restricted accessibility means that it still finds considerable use in certain situations and applications. Heavy construction, such as shipbuilding, and welding "in the field," away from many support services that would provide shielding gas, cooling water, and other necessities, rely on the SMAW process to a great extent. Although the SMAW process finds wide application for welding virtually all steels and many of the nonferrous alloys, it is primarily used to join steels. This family of materials includes low-carbon or mild steels, low-alloy steels, high-strength steels, quenched and tempered steels, high-alloy steels, stainless steels, and many of the cast irons. The SMAW process is also used to join nickel and its alloys and, to a lesser degree, copper and its alloys. It can be, but rarely is, used for welding aluminum.

In addition to joining metals, the SMAW process is frequently used for the protective surfacing of base metals. The surfacing deposit can be applied for the purpose of corrosion control or wear resistance (hard surfacing). Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Electrodes The electrodes used in the SMAW process have many different compositions of core wire and a wide variety of fluxcovering types and weights. Standard electrode diameters of the core wire range from 1.6 to 8 mm (

1 5 to in.). 16 16

Electrode length usually ranges from 230 to 455 mm (9 to 18 in.); the shorter lengths are associated with the smallerdiameter electrodes. A bare, uncoated end of the electrode (the grip end) is provided for making electrical contact in the electrode holder. The coating on the electrode has numerous functions. It provides: •

• • • • •

GAS (NORMALLY, CARBON DIOXIDE), FROM THE DECOMPOSITION OF CERTAIN COATING INGREDIENTS TO SHIELD THE ARC AND WELD ZONE FROM THE ATMOSPHERE DEOXIDIZERS, FOR SCAVENGING AND PURIFYING THE DEPOSITED WELD METAL SLAG FORMERS, TO PROTECT THE DEPOSITED WELD METAL FROM ATMOSPHERIC OXIDATION AND TO HELP SHAPE THE WELD BEAD IONIZING ELEMENTS, TO MAKE THE ARC MORE STABLE AND TO OPERATE WITH ALTERNATING CURRENT ALLOYING ELEMENTS, TO PROVIDE SPECIAL CHARACTERISTICS TO THE WELD DEPOSIT IRON POWDER, IN CERTAIN ELECTRODES, TO INCREASE PRODUCTIVITY FOR WELDING FERROUS METALS

The American Welding Society (AWS) has established a system for identifying and classifying the different types of welding electrodes. All SMAW electrodes have the prefix letter E to indicate welding electrode. The symbols that follow the prefix are based on criteria that best describe the welding capabilities of the electrode metal. These criteria include chemical composition of the deposited weld metal, weld-metal mechanical properties, certain process parameters, or combinations of all factors. Mild and Low-Alloy Steel-Covered Electrodes. The prefix used to identify these electrodes is followed by a number series that indicates minimum strength level, position capability, and type of covering and welding current. Table 2 explains how the number series is used in AWS A5.1, the specification for carbon steel electrodes for shielded metal arc welding and AWS A5.5, the specification for low-alloy steel electrodes. The first two digits after the E in the E6010 electrode designate a tensile strength of at least 430 MPa (62 ksi) for the deposited metal in the as-welded condition. The third digit indicates the position in which satisfactory welds can be made with the electrode. Thus, the 1 in E6010, for example, means that the electrode is satisfactory for use in all positions (flat, vertical, horizontal, and overhead). The 2 in E6020 indicates that the electrode is suitable for the flat position and horizontal fillets. The last digit or last two digits, taken together, indicate the applicable current type to be used and the type of covering on the electrode.

TABLE 2 CARBON AND LOW-ALLOY STEEL COVERED ELECTRODE IDENTIFICATION SYSTEM

AWS MINIMUM TENSILE MINIMUM YIELD MINIMUM STRENGTH CLASSIFICATION(A) STRENGTH ELONGATION, % MPA KSI MPA KSI

E70XX E80XX E90XX E100XX E110XX E120XX CLASSIFICATION(B) EXXIX EXX2X EXX4X

480-500 550 620 690 760 830 FLAT POSITION YES YES YES

CLASSIFICATION(C)

CURRENT

ARC

EXX10 EXXX1

DCEP AC AND DCEP AC AND DCEN AC AND DC AC AND DC DCEP AC OR DCEP AC OR DCEP AC OR DC AC OR DC AC OR DC AC OR

EXXX2 EXXX3 EXXX4 EXXX5 EXXX6 EXXX8 EXX20 EXX24 EXX27 EXX28

70-72 390-420 80 460-550 90 530-620 100 600 110 670-760 120 740-830 HORIZONTAL POSITION YES FILLET YES PENETRATIONN

57-60 67-80 77-90 87 97-110 107-120 VERTICAL POSITION YES NO DOWN

17-25 16-24 14-24 13-20 15-20 14-18 OVERHEAD POSITION YES NO YES

COVERING/SLAG

APPROXIMATE IRON POWDER(D), %

DIGGING DEEP DIGGING DEEP

CELLULOSE/SODIUM CELLULOSE/POTASSIUM

0-10 0

MEDIUM

MEDIUM

RUTILE/SODIUM

0-10

LIGHT

LIGHT

RUTILE/POTASSIUM

0-10

LIGHT

LIGHT

RUTILE/IRON POWDER

25-40

MEDIUM

MEDIUM

0

MEDIUM

MEDIUM

MEDIUM

MEDIUM

MEDIUM

MEDIUM

LOW HYDROGEN/SODIUM LOW HYDROGEN/POTASSIUM LOW HYDROGEN/IRON POWDER IRON OXIDE/SODIUM

LIGHT

LIGHT

RUTILE/IRON POWDER

50

MEDIUM

MEDIUM

MEDIUM

MEDIUM

IRON OXIDE/IRON 50 POWDER LOW HYDROGEN/IRON 50

0 25-40 0

(A) FIRST TWO OR THREE DIGITS INDICATE TENSILE STRENGTH IN UNITS OF KSI AND OTHER MECHANICAL PROPERTIES (MECHANICAL PROPERTY REQUIREMENTS VARY WITHIN EACH CLASSIFICATION). (B) THIRD (OR) FOURTH (SECOND TO LAST) DIGIT INDICATES THE WELDING POSITION THAT CAN BE USED. (C) LAST DIGIT INDICATES USABILITY OF THE ELECTRODE. (D) IRON POWDER PERCENTAGE BASED ON WEIGHT OF THE COVERING Stainless Steel Covered Electrodes. The three-digit number that follows the prefix E indicates the chemical composition. In addition, letters or numbers can be used to indicate composition modifications or position usability. The specification AWS A5.4 identifies and classifies covered corrosion-resisting chromium and chromium-nickel steel welding electrodes. Nickel and Copper Alloys. The designations for nonferrous product classifications, such as nickel and nickel alloys in

AWS A5.11 and copper and copper alloys in AWS A5.6, follow the prefix with a list of chemical element abbreviations that are significant in identifying product composition, such as ENiCu, ENiCrFe, ECuSi, and ECuNi.

Surfacing Welding Electrodes. The designations for these products are contained in specifications AWS A5.13 and

A5.21. They are very similar to the system used to identify nonferrous electrodes. Aluminum and Aluminum Alloys. The specification for aluminum and aluminum alloy arc welding electrodes, AWS

A5.3, uses the E prefix to indicate a covered electrode, followed by a series of numbers that identify the chemical composition that is equivalent to Aluminum Association alloy designations (for example, E1100, E3003, and E4043). Suffix symbols are used in various classifications. The AWS A5.5 specification for low-alloy filler metals uses suffixes

such as -A1, -B2, -B2L, and -C1 to indicate chemical compositions. Table 3 identifies the weld deposit chemical composition associated with a number of suffixes found on low-alloy electrodes. The classifications for nonferrous products in the AWS A5.6 specification for copper alloys and in the AWS A5.11 specification for nickel alloys list a letter or number suffix that indicates position in a series of similar alloy groupings. A similar suffix pattern is also used in the AWS A5.13 and AWS A5.21 specifications for surfacing welding electrodes. Covered stainless steel electrodes employ a number, -15 or -16, as a suffix to identify usability. The -15 suffix indicates that the electrode is designed for all-position operation using DCEP electrical current. The -16 suffix indicates all-position operation with either ac or DCEP.

TABLE 3 SUFFIX SYMBOLS AND CORRESPONDING COMPOSITIONS FOR LOW-ALLOY STEEL COVERED ELECTRODES

SUFFIX(A) COMPOSITION, % C Mn A1 0.12 0.6-1.0(B) B1 0.12 0.90 B2L 0.05 0.90 B2 0.12 0.90 B3L 0.05 0.90 B3 0.12 0.90 B4L 0.05 0.90 B5 0.07-0.15 0.40-0.70 C1 0.12 1.20 C2 0.12 1.20 C3 0.12 0.40-1.25 D1 0.12 1.25-1.75 D2 0.15 1.65-2.00 G ... 1.0 MIN (C) M 0.10 0.6-2.25(B)

Si 0.40-0.80(B) 0.60-0.80(B) 0.8-1.0(B) 0.60-0.80(B) 0.8-1.0(B) 0.60-0.80(B) 1.00 0.30-0.60 0.6-0.8(B) 0.6-0.8(B) 0.80 0.6-0.8(B) 0.6-0.8(B) 0.80 MIN 0.6-0.8(B)

Ni ... ... ... ... ... ... ... ... 1.00-2.75 3.00-3.75 0.80-1.10 ... ... 0.50 MIN 1.4-2.5(B)

Cr ... 0.40-0.65 1.00-1.50 1.00-1.50 2.00-2.50 2.00-2.50 1.75-2.25 0.50-0.60 ... ... 0.15 ... ... 0.30 MIN 0.15-1.5(B)

Mo 0.40-0.65 0.40-0.65 0.40-0.65 0.40-0.65 0.90-1.20 0.90-1.20 0.40-0.65 1.00-1.25 ... ... 0.35 0.25-0.45 0.25-0.45 0.20 MIN 0.25-0.55(B)

V ... ... ... ... ... ... 0.05 ... ... 0.05 ... ... 0.10 MIN 0.05

(A) THE SUFFIX INDICATES THE CHEMICAL COMPOSITION OF THE WELD-METAL DEPOSIT. (B) AMOUNT DEPENDS ON ELECTRODE CLASSIFICATION. SINGLE VALUES INDICATE MAXIMUM, CHECK AWS A5.5 FOR THE DIFFERENT ELECTRODE CLASSES. (C) THERE ARE SEVERAL DIFFERENT M CLASSES; M CLASSIFICATIONS ARE INTENDED TO CONFORM TO MILITARY SPECIFICATIONS. Deposition Rates. The melting rate of the electrode is directly related to the welding current. The current density in the

electrode increases with higher current, which increases the melting rate, which, in turn, increases the deposition rate. The electrode coating also affects deposition rate. The iron powder types are designed to have higher deposition rates and, therefore, greater productivity. Figure 4 shows the expected deposition rate versus amperage for various electrodes at a 100% duty cycle. (The actual deposition rate will be considerably less. Deposition rate is a function of the duty cycle, which is affected by the time spent changing the electrodes, cleaning slag off of the weld, etc.) Electrode size and, therefore, usable current range are determined by the base-metal thickness, welding position, welder skill level, and joint details.

FIG. 4 DEPOSITION RATE VERSUS AMPERES FOR VARIOUS ELECTRODES

Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Weld Schedules and Procedures Welding schedules are tables of operating parameters that will provide high-quality welds under normal conditions.

Strict welding schedules are not as important for the manual SMAW process as they are for semiautomatic and automatic welding for several reasons. First, in a manual welding process, the welder controls conditions by arc manipulation, which achieves better control than any of the other arc welding processes. The welder also directly controls the arc voltage and travel speed and, indirectly, the welding current. Second, meter readings are rarely used in the SMAW process for the duplication of jobs. It is generally considered that the recommended welding current ranges given in Table 4 for the different types of electrodes are sufficient for most operations. The settings provide a good starting point when first welding on a new application, although they are not necessarily the only welding settings that can be used under every condition. For example, for high-production work, the current settings could be increased considerably over those shown. Factors such as weld appearance, welding position, and welder skill also allow variations from the settings.

TABLE 4 TYPICAL AMPERAGE RANGES FOR SELECTED SMAW ELECTRODES ELECTRODE DIAMETER mm in.

E6010 AND E6011

E6012

E6013

E6020

E6022

E6027 AND E7027

E7014

E7015, E7016, AND E7016-1

E7018 AND E7018-1

E70241, E7024, AND E7028

E7048

1.6

...

20-40

20-40

...

...

...

...

...

...

...

...

...

25-60

25-60

...

...

...

...

...

...

...

...

40-80

35-85

45-90

...

...

...

80140 110190 140240

80130 105180 150230

100150 130190 175250

110160 140190 170400

125185 160240 210300

65110 100150 140200 180255

70100 115165 150220 200275

100145 140190 180250 230305

...

75125 110170 140215

80125 110160 150210 200275

2.0 2.4(A) 3.2 4.0 4.8

1 16 5 64 3 32 1 8 5 32 3 16

80140 150220 210270

5.6 6.4 8.0(A)

7 32 1 4 5 16

170250 210320 275425

200320 250400 300500

210300 250350 320430

225310 275375 340450

370520 ... ...

250350 300420 375475

260340 330415 390500

240320 300390 375475

260340 315400 375470

275365 335430 400525

... ... ...

(A)

(A) THESE DIAMETERS ARE NOT MANUFACTURED IN THE E7028 CLASSIFICATION. Welder Training. The SMAW process generally requires a high degree of welder skill to consistently produce quality

welds. As a result, many training programs emphasize the SMAW process because of the arc manipulation skills developed by the welder. This acquired skill level makes the training on other processes much easier. The exact content of a training program will vary, depending on the specific application of the process. The complexity of the parts to be welded and the governing codes or specifications involved also dictate the length of the training program. For example, because a pipe welder would need more skill than a tack welder, the length of his training program would be greater. The job title of arc welder (DOT 810.384-014) describes a person whose responsibility is to weld together many components. This job includes setting up the machine and part to be welded, striking the arc and guiding it along the joint, and performing such duties as chipping, grinding, and slag removal. The welder should be able to: • • •

WELD IN ALL POSITIONS PASS EMPLOYER PERFORMANCE TESTS MEET CERTIFICATION STANDARDS OF GOVERNMENTAL AGENCIES OR PROFESSIONAL AND TECHNICAL ASSOCIATIONS

A tack welder (DOT 810.684-010) makes short beads at specific points to hold the parts in place for final welding. The tack welder also performs the duties of fitter helper. A combination welder (DOT 819-384-010) welds metal parts together to either fabricate or repair assemblies. The combination welder is able to create gas welds and electric arc welds, and to perform flame-cutting operations. The welder portion of the pipefitter description (DOT 862.381-018) must weld the pipe joint after it has been assembled and tacked in place. Procedures. There is a definite relationship between the welding current, the size of the welding electrodes, and the

welding position. These parameters must be selected to ensure that the welder has the molten weld-metal pool under complete control at all times. If the pool becomes too large, then it becomes unmanageable, perhaps allowing molten metal to run out, particularly during out-of-position welding. The welder should maintain the steady frying and crackling sound that comes with the use of correct procedures. The shape of the molten pool and the movement of the metal at the rear of the pool both serve as guides in checking weld quality. The ripples produced on the bead should be uniform and have good side tie-in with no undercut or excessive reinforcement. The factors described below are essential for maintaining high-quality welding. Correct Electrode Type. It is important to select the proper electrode for each job. The selection should be based on the type of base metal, expected service, and mechanical properties required. Correct Electrode Size. The choice of electrode size should depend on the type of electrode, welding position, joint

preparation, base-metal thickness, and welder skill. Correct Current. If the current is too high, then the electrode melts too fast and the molten pool becomes large,

irregular, and difficult to control. Current that is too low will not provide enough heat to melt the base metal, causing the molten pool to be sluggish, with a high, irregular, ropey weld bead. It should also be noted that the electrode has inherent

current limits. If the current is too high, then the core wire overheats and the coating cracks. If the current is too low, then there is insufficient heat to maintain the arc and form the protective gas shield. Correct Arc Length. If an arc is too long, then the metal melts off the electrode in large globules that wobble from side to side, resulting in a wide and irregular weld bead with considerable spatter, and, possibly, porosity and mechanical property degradation. If the arc is too short, then it has insufficient heat to melt the base metal and electrode, which often results in the electrode sticking to the work. Correct Travel Speed. A speed that is too fast allows the weld pool to freeze before impurities and gases can escape,

and the bead will be narrow and inadequate in size. When the speed is too slow, the metal piles up and the bead is larger than required. Correct Electrode Angle. In fillet welding and deep groove welding, the electrode angle is particularly important.

When making a fillet weld, the electrode should be held so that it bisects the angle between the plates and is perpendicular to the line of the weld. When undercut occurs in the vertical member, the angle should be lowered and the arc directed toward the vertical member. Correct Arc Manipulation. When weaving is used, the width of the weave and the pause at the ends of the weave become important. The welder must pause at each end of the weave to allow adequate fill buildup and fusion to occur. The welder should also quickly move across the center of the weld, because heating is more concentrated in the center than at the edges. Breaking the Arc. Before an arc is broken, it is important to know whether it will be reestablished with the next

electrode and the weld continued or whether it is the end of a weld pass. If welding is to continue, then the crater should remain and the arc quickly broken. If it is the end of a weld pass, then the arc should not be broken until the crater has filled. Correct Interbead Cleaning. Proper interbead cleaning to remove slag and any spatter is essential to the production of

high-quality welds. Proper cleaning prevents slag inclusions, "lack of fusion" defects, and porosity. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Variations of the SMAW Process Gravity Welding. Gravity feed is considered to be an automatic method of applying the SMAW process. It utilizes a relatively low-cost mechanism that includes an electrode holder attached to a bracket, which slides down an inclined bar arranged along the line of weld. Special electrodes with a heavy coating are maintained in contact with the workpiece by the weight of the electrode holder and electrode. Once the process is started, it continues automatically until the electrode has burned to a short stub, whereupon the bracket and electrode holder are automatically kicked up to break the arc. One welder can operate several gravity feeders at the same time (Fig. 5). This increases productivity, reduces welder fatigue, and requires less-skilled welders, all of which result in substantial savings in welder labor costs. Table 5 compares the deposition rate, based on pounds per hour, when using one electrode manually versus two, three, four, or five gravity feeders.

TABLE 5 COMPARISON OF DEPOSITION RATES FOR CONVENTIONAL SMAW AND MULTI-ARC GRAVITY-FED WELDING Data are for making 7.9 mm (

5 in.) fillet welds using E6027 electrodes. 16

METHOD MANUAL, ONE ARC GRAVITY, TWO ARCS GRAVITY, THREE ARCS

DEPOSITION RATE, LB/H 9 17 26

GRAVITY, FOUR ARCS GRAVITY, FIVE ARCS

34 43

FIG. 5 GRAVITY FEEDERS BEING USED ON SHIP SUBASSEMBLIES

Although gravity-fed SMAW was investigated in the United States, England, and the Scandinavian countries in the late 1940s and early 1950s, credit must be given to the Japanese shipbuilders for perfecting and utilizing the process on a large scale in the early 1960s. The gravity welding process is being used in shipyards, railroad car shops, and barge yards throughout the world. It has reasonable acceptance in applications where large amounts of horizontal fillet welds must be made in a relatively small area. Firecracker welding is an automatic method by which shielded metal arc welds are made using a long electrode with an electrically nonconductive coating. Human involvement is not required after the arc is initiated. A firecracker weld is generally positioned in the flat position. The welding electrode is placed in the joint and a retaining bar is placed over it. The arc is started by shorting the end of the electrode to the workpiece. The arc length is controlled by the coating thickness. As the arc travels along the stationary electrode, the electrode melts and makes a deposit on the metal immediately underneath. Once the arc is started, the process proceeds to completion automatically. Electrodes that are up to 1 m (39 in.) long and have a core diameter of 5, 6, or 8 mm (0.20, 0.24, or 0.32 in.) have been used. Both alternating and direct current have been applied, and the former may be preferred, because of arc blow problems associated with direct current. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Special Applications of the SMAW Process Underwater welding began during World War I when the British naval force used it to make temporary repairs of

leaking rivets on ship hulls. The introduction of covered electrodes enabled succesful underwater welding and the production of welds having approximately 80% of the strength and 40% of the ductility of similar welds made in air. Because of the somewhat diminished weld properties, this SMAW application is generally restricted to salvage operations or underwater repair work. Underwater welding can be subdivided into two major categories: welding in a wet environment and welding in a dry environment. Wet Welding. The relatively poor quality of welds made in a wet environment is due primarily to the problem of heat

transfer, welder visibility, and the presence of hydrogen in the arc atmosphere during the welding operation. When the base metal and the arc area are surrounded entirely by water, there is no temperature or heat buildup of the base metal at the weld, nor is preheating possible. This creates a weld-metal quench effect, which traps damaging amounts of hydrogen and also produces a weld solidification structure with reduced toughness and ductility. Both conditions contribute to the weld-metal cracking problems experienced when welding steels underwater. Another disadvantage is the restricted visibility, which is due to the equipment and the existing local contaminants in the water, as well as those generated by the welding arc. Under the most ideal conditions, welds produced in wet environments using covered electrodes are

marginal, at best. They can be placed in service for short periods of time under reduced operating conditions, but should be replaced with quality welds as quickly as possible. The covered electrodes used for wet welding must be waterproofed prior to underwater use. This can be done by wrapping them with waterproof tape or by dipping them in special sodium silicate mixes and allowing them to dry. Dry Welding. The dry environment enables the production of high-quality welds that meet all code quality requirements. The SMAW process is not very popular for welding in the dry environment, because large amounts of smoke and fumes are produced. An extensive air moving, filtering, and refrigeration system must be employed when the SMAW process is used, because a dry-environment area will quickly fill with the welding fumes, making it impossible for the welder to see that weld area and to function. For this reason, the gas-tungsten arc welding and gas-metal arc welding processes have broader use in dry welding applications.

The article "Underwater Welding" in this Volume explores the topic in greater detail. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Safety Considerations The SMAW process, like all open arc welding processes, has a number of potential hazards. To alert welders about these safety concerns, all SMAW electrode containers carry a warning label that identifies the three most common safety hazards in these terms: •

•

•

ELECTRIC SHOCK CAN KILLDO NOT PERMIT ELECTRICALLY LIVE PARTS OR ELECTRODES TO CONTACT SKIN . . . OR YOUR CLOTHING OR GLOVES IF THEY ARE WETINSULATE YOURSELF FROM WORK AND GROUND FUMES AND GASES CAN BE DANGEROUS TO YOUR HEALTHKEEP FUMES AND GASES FROM YOUR BREATHING ZONE AND GENERAL AREAKEEP YOUR HEAD OUT OF FUMESUSE ENOUGH VENTILATION OR EXHAUST AT THE ARC, OR BOTH ARC RAYS CAN INJURE EYES AND BURN SKINWEAR CORRECT EYE, EAR, AND BODY PROTECTION

In addition to the general warnings on the container, further detailed information relating to safety is contained in ANSI/AWS Z49.1, "Safety in Welding and Cutting." The information described below is intended to expand on the general warnings, and should be particularly useful for SMAW welders. First, the filter plates installed within welding shields must be capable of stopping the harmful levels of infrared, ultraviolet, and visible light rays originating in the arc. Filter plates are now able to absorb 99% or more of the infrared and ultraviolet rays from the arc. The shade of the filter plate suggested for use with SMAW electrodes is given in Table 6.

TABLE 6 RECOMMENDED FILTER LENS SHADES USED IN SHIELDED METAL ARC WELDING

ELECTRODE DIAMETER LENS SHADE NO. mm in. 1 1.6 10 2.4 3.2

16 3 32 1 8

4.0 4.8 5.6 6.4 7.9 9.5

5 32 3 16 7 32 1 4 5 16 3 8

12

14

Source: ANSI/AWS Z49.1

Second, in addition to protecting himself, the welder must also be aware of others in the area who need protection, which can usually be provided by portable screens. The failure of those working around the arc to use adequate protection can result in eye burn, which is similar to sunburn and is extremely painful for a period of up to 48 h. Usually, eye burn does not permanently injure the eyes, but it can cause intense pain. A physician should be consulted in the case of severe arc burn, regardless of whether it involves the skin or the eyes. Third, the welding area must be adequately ventilated because of the heavy concentrations of smoke and fumes generated in the SMAW process. If welding is being performed in confined spaces with poor ventilation, such as in a tank, an external air supply in the form of a mask or special helmet may be required. In addition, a second person should be stationed at the tank manhole to provide any necessary assistance. Special ventilation is required when welding stainless steels or metals coated with copper, zinc, lead, or cadmium, because of the toxic nature of these fumes. Fourth, cables with frayed or cracked insulation and faulty or badly worn connections can cause electrical short circuits and electrical shocks to personnel. When it is necessary to weld in a damp or wet area, the welder should wear rubber boots and stand on a dry, insulated platform. Shielded Metal Arc Welding Raymond H. Juers, Naval Surface Warfare Center

Selected References

• H.B. CARY, MODERN WELDING TECHNOLOGY, 2ND ED., PRENTICE-HALL, 1989 • "FILLER METAL COMPARISON CHARTS," AWS FMC-89, AMERICAN WELDING SOCIETY • "SAFETY IN WELDING AND CUTTING," ANSI/AWS Z49.1, AMERICAN WELDING SOCIETY • "TECHNICAL GUIDE FOR SHIELDED METAL ARC WELDING," HOBART BROTHERS CO. Gas-Metal Arc Welding D.B. Holliday, Westinghouse Electric Corporation

Introduction GAS-METAL ARC WELDING (GMAW) is an arc welding process that joins metals together by heating them with an electric arc that is established between a consumable electrode (wire) and the workpiece. An externally supplied gas or gas mixture acts to shield the arc and molten weld pool. Although the basic GMAW concept was introduced in the 1920s, it was not commercially available until 1948. At first, it was considered to be fundamentally a high-current-density, small-diameter, bare-metal electrode process using an inert

gas for arc shielding. Its primary application was aluminum welding. As a result, it became known as metal-inert gas (MIG) welding, which is still common nomenclature. Subsequent process developments included operation at low current densities and pulsed direct current, application to a broader range of materials, and the use of reactive gases (particularly carbon dioxide) and gas mixtures. The latter development, in which both inert and reactive gases are used, led to the formal acceptance of the term gas-metal arc welding. The GMAW process can be operated in semi-automatic and automatic modes. All commercially important metals, such as carbon steel, high-strength low-alloy steel, stainless steel, aluminum, copper, and nickel alloys can be welded in all positions by this process if appropriate shielding gases, electrodes, and welding parameters are chosen. Advantages. The applications of the process are dictated by its advantages, the most important of which are: • • • • • •

•

•

ELECTRODE LENGTH DOES NOT FACE THE RESTRICTIONS ENCOUNTERED WITH SHIELDED-METAL ARC WELDING (SMAW). WELDING CAN BE ACCOMPLISHED IN ALL POSITIONS, WHEN THE PROPER PARAMETERS ARE USED, A FEATURE NOT FOUND IN SUBMERGED ARC WELDING. WELDING SPEEDS ARE HIGHER THAN THOSE OF THE SMAW PROCESS. DEPOSITION RATES ARE SIGNIFICANTLY HIGHER THAN THOSE OBTAINED BY THE SMAW PROCESS. CONTINUOUS WIRE FEED ENABLES LONG WELDS TO BE DEPOSITED WITHOUT STOPS AND STARTS. PENETRATION THAT IS DEEPER THAN THAT OF THE SMAW PROCESS IS POSSIBLE, WHICH MAY PERMIT THE USE OF SMALLER-SIZED FILLET WELDS FOR EQUIVALENT STRENGTHS. LESS OPERATOR SKILL IS REQUIRED THAN FOR OTHER CONVENTIONAL PROCESSES, BECAUSE THE ARC LENGTH IS MAINTAINED CONSTANT WITH REASONABLE VARIATIONS IN THE DISTANCE BETWEEN THE CONTACT TIP AND THE WORKPIECE. MINIMAL POSTWELD CLEANING IS REQUIRED BECAUSE OF THE ABSENCE OF A HEAVY SLAG.

These advantages make the process particularly well suited to high-production and automated welding applications. With the advent of robotics, gas-metal arc welding has become the predominant process choice. Limitations. The GMAW process, like any welding process, has certain limitations that restrict its use: • •

THE WELDING EQUIPMENT IS MORE COMPLEX, USUALLY MORE COSTLY, AND LESS PORTABLE THAN SMAW EQUIPMENT. THE PROCESS IS MORE DIFFICULT TO APPLY IN HARD-TO-REACH PLACES BECAUSE THE WELDING GUN IS LARGER THAN A SMAW HOLDER AND MUST BE HELD CLOSE TO THE JOINT (WITHIN 10 TO 19 MM, OR

•

•

3 8

TO

3 4

IN.) TO ENSURE THAT THE WELD METAL IS

PROPERLY SHIELDED. THE WELDING ARC MUST BE PROTECTED AGAINST AIR DRAFTS THAT CAN DISPERSE THE SHIELDING GAS, WHICH LIMITS OUTDOOR APPLICATIONS UNLESS PROTECTIVE SHIELDS ARE PLACED AROUND THE WELDING AREA. RELATIVELY HIGH LEVELS OF RADIATED HEAT AND ARC INTENSITY CAN HINDER OPERATOR ACCEPTANCE OF THE PROCESS.

Gas-Metal Arc Welding D.B. Holliday, Westinghouse Electric Corporation

Process Fundamentals Principles of Operation. In the GMAW process (Fig. 1), an arc is established between a continuously fed electrode of filler metal and the workpiece. After proper settings are made by the operator, the arc length is maintained at the set value, despite the reasonable changes that would be expected in the gun-to-work distance during normal operation. This automatic arc regulation is achieved in one of two ways. The most common method is to utilize a constant-speed (but adjustable) electrode feed unit with a variable-current (constant-voltage) power source. As the gun-to-work relationship changes, which instantaneously alters the arc length, the power source delivers either more current (if the arc length is decreased) or less current (if the arc length is increased). This change in current will cause a corresponding change in the electrode melt-off rate, thus maintaining the desired arc length.

FIG. 1 SCHEMATIC OF GMAW PROCESS

The second method of arc regulation utilizes a constant-current power source and a variable-speed, voltage-sensing electrode feeder. In this case, as the arc length changes, there is a corresponding change in the voltage across the arc. As this voltage change is detected, the speed of the electrode feed unit will change to provide either more or less electrode per unit of time. This method of regulation is usually limited to larger electrodes with lower feed speeds. Metal Transfer Mechanisms. The characteristics of the GMAW process are best described by reviewing the three basic means by which metal is transferred from the electrode to the work: short-circuiting transfer, globular transfer, or spray transfer. The type of transfer is determined by a number of factors, the most influential of which are: • • • • • •

MAGNITUDE AND TYPE OF WELDING CURRENT ELECTRODE DIAMETER ELECTRODE COMPOSITION ELECTRODE EXTENSION BEYOND THE CONTACT TIP OR TUBE SHIELDING GAS POWER SUPPLY OUTPUT

Short-circuiting transfer encompasses the lowest range of welding currents and electrode diameters associated with the

GMAW process. This type of transfer produces a small, fast-freezing weld pool that is generally suited for joining thin sections, for out-of-position welding, and for bridging of large root openings. Metal is transferred from the electrode to the workpiece only during a period when the electrode is in contact with the weld pool, and there is no metal transfer across the arc gap (Fig. 2).

FIG. 2 TRANSFER MODES IN GMAW PROCESS

The electrode contacts the molten weld pool at a steady rate that can range from 20 to over 200 times per second. As the wire touches the weld metal, the current increases and the liquid metal at the wire tip is pinched off, initiating an arc. The rate of current increase must be high enough to heat the electrode and promote metal transfer, yet low enough to minimize spatter caused by violent separation of the molten drop. The rate of current increase is controlled by adjusting the power source inductance. The optimum setting depends on the electrical resistance of the welding circuit and the melting temperature of the electrode. When the arc is initiated, the wire melts at the tip as it is fed forward toward the next short circuit. The open-circuit voltage of the power source must be low enough so that the drop of molten metal cannot transfer until it contacts the weld metal. Because metal transfer only occurs during short circuiting, the shielding gas has very little effect on the transfer itself. However, the gas does influence the operating characteristics of the arc and the base-metal penetration. The use of carbon dioxide generally produces high spatter levels, when compared with inert gases, but it allows deeper penetration when welding steels. To achieve a good compromise between spatter and penetration, mixtures of carbon dioxide and argon are often used. With nonferrous metals, argon-helium mixtures are used to achieve this compromise. Globular Transfer. With a positive electrode, globular transfer takes place when the current density is relatively low,

regardless of the type of shielding gas. However, the use of carbon dioxide or helium results in this type of transfer at all usable welding currents. Globular transfer is characterized by a drop size with a diameter that is greater than that of the electrode. This large drop is easily acted upon by gravity, which limits successful transfer to the flat position.

At average currents that are slightly higher than those used in short-circuiting transfer, axially directed globular transfer can be achieved in a substantially inert gas shield. However, if the arc length is too short, then the enlarging drop can short to the workpiece, become superheated, and disintegrate, producing considerable spatter. Therefore, the arc length must be long enough to ensure that the drop detaches before it contacts the weld pool. However, when higher voltage values are used, the weld is likely to be unacceptable, because of a lack of fusion, insufficient penetration, and excessive reinforcement. This limits the use of this transfer mode to very few production applications. Carbon dioxide shielding produces a randomly directed globular transfer when the welding current and voltage values are significantly higher than the range used for short-circuiting transfer. Although severe spatter conditions result when conventional techniques are used, carbon dioxide is still the most commonly used shielding gas for welding mild steel when the quality requirements are not too rigorous. The spatter problem is controlled by "burying" the arc below the weld/base-metal surface. The resulting arc forces are adequate enough to produce a depression that contains the spatter. This technique requires relatively high currents and results in very deep penetration. Good operator setup skills are required. However, poor wetting action can result in an excessive weld reinforcement. Spray Transfer. A very stable, spatter-free "spray" transfer mode can be produced when argon-rich shielding is used. This type of transfer requires the use of direct current with the electrode positive and a current level that is above a critical value called the "transition current." Below this current level, transfer occurs in the globular mode at the rate of a few drops per second. At values above the transition current, transfer occurs in the form of very small drops that are formed and detached at the rate of hundreds per second and are accelerated axially across the arc gap.

The transition current is proportional to the electrode diameter, and, to a lesser extent, to the electrode extension. It also has a direct relationship to the filler metal melting temperature. Transition currents for various materials and electrode diameters are shown in Table 1.

TABLE 1 GLOBULAR-TO-SPRAY TRANSITION CURRENTS FOR SELECTED ELECTRODES

WIRE ELECTRODE TYPE LOW-CARBON STEEL

STAINLESS STEEL

SHIELDING WIRE DIAMETER GAS mm in. 98AR-2O2 0.58 0.023 0.76 0.030 0.89 0.035 1.14 0.045 1.57 0.062 95AR-5O2 0.89 0.035 1.14 0.045 1.57 0.062 92AR-8CO2 0.89 0.035 1.14 0.045 1.57 0.062 85AR-15CO2 0.89 0.035 1.14 0.045 1.57 0.062 80AR-20CO2 0.89 0.035 1.14 0.045 1.57 0.062 99AR-1O2 0.89 0.035 1.14 0.045 1.57 0.062 AR-HE-CO2 0.89 0.035 1.14 0.045 1.57 0.062 AR-H2-CO2 0.89 0.035

SPRAY ARC CURRENT, A 135 150 165 220 275 155 200 265 175 225 290 180 240 295 195 255 345 150 195 265 160 205 280 145

ALUMINUM

ARGON

DEOXIDIZED COPPER ARGON

SILICON BRONZE

ARGON

1.14 1.57 0.76 1.19 1.57 0.89 1.14 1.57 0.89 1.14 1.57

0.045 0.062 0.030 0.047 0.062 0.035 0.045 0.062 0.035 0.045 0.062

185 255 95 135 180 180 210 310 165 205 270

Source: Union Carbide Industrial Gases

The spray transfer mode results in a highly directed stream of discrete drops that are accelerated by arc forces to velocities that overcome the effects of gravity. This enables the process to be used in any position, under certain conditions. Because the drops are separated, short circuits do not occur, and the spatter level is negligible, if not totally eliminated. Another characteristic of spray transfer is the "finger" penetration pattern that it produces directly below the electrode tip. Although the penetration can be deep, it can be affected by magnetic fields that must be controlled to ensure that it is always located at the center of the weld penetration profile. Otherwise, a lack of fusion and an irregular bead surface profile can result. The spray transfer mode can be used to weld almost any metal or alloy, because of the inert characteristics of the argon shield. Sometimes, thickness can be a factor, because of the relatively high current levels required. The resultant arc forces can cut through, rather than weld, thin sheets. In addition, high deposition rates can result in a weld pool size that cannot be supported by surface tension in the vertical and overhead positions. However, the thickness and position limitations of spray transfer have been largely overcome by specially designed power supplies. These machines produce carefully controlled current outputs that "pulse" the welding current from levels below the transition current to levels above it. Figure 3 shows the two levels of current provided by these machines. One is a constant, low-background current that sustains the arc without providing enough energy to cause the formation of drops on the wire tip. The other is a superimposed pulsing current with an amplitude that is greater than the transition current necessary for spray transfer. During this pulse, one or more drops are formed and transferred. The frequency and amplitude of the pulses control the energy level of the arc and, therefore, the rate at which the wire melts. By reducing the arc energy and the wire melting rate, it is possible to retain many of the desirable features of spray transfer while joining sheet metals and welding thick metals in all positions.

FIG. 3 CHARACTERISTIC CURRENT WAVEFORM FOR A "PULSING" POWER SUPPLY

Many variations of such machines are available. The simplest provide a single frequency of pulsing (60 or 120 pulses/s) and independent control of the background and pulsing current levels. Synergic machines, which are sophisticated, automatically provide the optimum combination of background and pulsing current levels for any given setting of wire feed speed. Normally, these settings are specific to an electrode/shielding gas combination and must be changed or reprogrammed when the combination is changed. Process Variables. The important variables of the GMAW process that affect weld penetration, bead geometry, and

overall weld quality are: • • • • • • •

WELDING CURRENT (ELECTRODE FEED SPEED) POLARITY ARC VOLTAGE (ARC LENGTH) TRAVEL SPEED ELECTRODE EXTENSION ELECTRODE ORIENTATION (GUN ANGLE) ELECTRODE DIAMETER

Knowledge and control of these variables are essential to consistently produce welds of satisfactory quality. Because they are not completely independent of one another, changing one variable generally requires changing one or more of the others to produce the desired results. The effects of these variables on deposit attributes are shown in Table 2.

TABLE 2 EFFECT OF CHANGES IN PROCESS VARIABLES ON WELD ATTRIBUTES WELDING VARIABLES TO CHANGE

CURRENT AND WIRE FEED SPEED VOLTAGE TRAVEL SPEED ELECTRODE EXTENSION WIRE DIAMETER SHIELD GAS % GUN ANGLE

DESIRED CHANGES PENETRATION DEPOSITION RATE

BEAD SIZE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

LITTLE EFFECT

LITTLE EFFECT

NO EFFECT NO EFFECT DECREASE

NO EFFECT NO EFFECT INCREASE

LITTLE EFFECT LITTLE EFFECT INCREASE(A)

LITTLE EFFECT LITTLE EFFECT DECREASE(A)

LITTLE EFFECT DECREASE

LITTLE EFFECT INCREASE

INCREASE

DECREASE

DECREASE

INCREASE

INCREASE

DECREASE

DECREASE

INCREASE

DECREASE

INCREASE

DECREASE

INCREASE

INCREASE

DECREASE

LITTLE EFFECT DECREASE

PUSH

LITTLE EFFECT LITTLE EFFECT

LITTLE EFFECT LITTLE EFFECT LITTLE EFFECT

LITTLE EFFECT INCREASE

DRAG

LITTLE EFFECT LITTLE EFFECT

LITTLE EFFECT LITTLE EFFECT LITTLE EFFECT

PUSH

DRAG

BEAD WIDTH

(A) WILL RESULT IN DESIRED CHANGE IF CURRENT LEVELS ARE MAINTAINED BY ADJUSTMENT OF WIRE FEED SPEED Considerable skill and experience are necessary to select the optimal combination for each application. This selection is further complicated by the fact that the optimal settings for the variables are also affected by the type of base metal, the electrode composition, the welding position, quality requirements, and the number of completed weldments required. Thus, no single set of parameters provides optimal results in every case. Welding Current. As the electrode feed speed is varied, the welding current varies in a like manner when a constantvoltage power source is used. This occurs because the current output of the power source varies dramatically with the slight changes in the arc voltage (arc length) that result when changes are made in the electrode feed speed. When all

other variables are held constant, an increase in welding current results in an increase in the depth and width of penetration, deposition rate, and weld bead size. Polarity is the term used to described the electrical connection of the welding gun in relation to the terminals of a direct-

current (dc) power source. When the gun power lead is connected to the positive terminal, the polarity is designated as direct current, electrode positive (DCEP). Alternatively, a connection to the negative terminal is designated as direct current, electrode negative (DCEN). The vast majority of GMAW applications utilize DCEP, because it provides for a stable arc, low spatter, a good weld bead profile, and the greatest depth of penetration. Arc voltage and arc length are related terms that are often used interchangeably. However, they are different. Arc voltage

is an approximate means of stating the physical arc length in electrical terms. The same physical arc length, however, could yield different arc voltage readings, depending on factors such as shielding gas, current, and electrode extension. When all variables are held constant, a reliable relationship exists between the two: an increase in voltage setting will result in longer arc length. Although the arc length is the variable of interest and the one that should be controlled, arc voltage is more easily monitored. Because of this fact, and because the arc voltage is normally required to be specified in welding procedures, it is the term that is more commonly used. From any specific value of arc voltage, an increase tends to flatten the weld bead and increase the width of the fusion zone. Excessively high voltage can cause porosity, spatter, and undercut. A reduction in voltage results in a narrower weld bead with a higher crown. Travel speed is the linear rate at which the arc is moved along the weld joint. When all other conditions are held

constant, weld penetration is a maximum at an intermediate travel speed. When travel speed is decreased, the filler metal deposition per unit length increases. At very slow speeds, the welding arc impinges on the molten weld pool, rather than the base metal, thereby reducing the effective penetration. As the travel speed is increased, the thermal energy transmitted to the base metal from the arc increases, because the arc acts more directly on the base metal. However, further increases in travel speed impart less thermal energy to the base metal. Thus, melting of the base metal first increases and then decreases with increasing travel speed. As travel speed is increased further, there is a tendency toward undercutting along the edges of the weld bead, because there is insufficient deposition of filler metal to fill the path melted by the arc. Electrode orientation is described in two ways: by the relationship of the electrode axis with respect to the direction of

travel (the travel angle) and by the angle between the electrode axis and the adjacent work surface (work angle). When the electrode points in a direction opposite to the travel direction, it results in a trail angle and is known as the backhand welding technique. When the electrode points in the direction of travel, it results in a lead angle and is called the forehand welding technique. For all positions, a trailing travel angle that ranges from 5 to 15° (from perpendicular) provides a weld with maximum penetration and a narrow, convex surface configuration. It also provides for maximum shielding of the molten weld pool. However, the common technique utilizes a leading travel angle, which provides better visibility for the operator and a weld with a flatter surface profile. For some materials, such as aluminum, a leading angle is preferred, because it provides a "cleaning action" ahead of the molten weld metal, which promotes wetting and reduces base-material oxidation. When producing fillet welds in the horizontal position, the work angle should be about 45° to the vertical member. The electrode extension is the distance between the last point of electrical contact (usually the gun contact tip or tube)

and the end of the electrode. An increase in the amount of this extension causes an increase in electrical resistance. This, in turn, generates additional heat in the electrode, which contributes to greater electrode melting rates. Without an increase in arc voltage, the additional metal will be deposited as a narrow, high-crowned weld bead. The optimum electrode extension generally ranges from 6.4 to 13 mm ( (

1 1 to in.) for short-circuiting transfer and from 13 to 25 mm 4 2

1 to 1 in.) for spray and globular transfers. 2

The electrode diameter influences the weld bead configuration. A larger electrode requires a higher minimum current

than a smaller electrode does to achieve the same metal transfer characteristics.

Higher currents, in turn, produce additional electrode melting and larger, more-fluid weld deposits. Higher currents also result in higher deposition rates and greater penetration, but may prevent the use of some electrodes in the vertical and overhead positions. Gas-Metal Arc Welding D.B. Holliday, Westinghouse Electric Corporation

Equipment The basic equipment for a typical GMAW installation is shown in Fig. 1. The major components are discussed below. A welding gun provides electrical current to the electrode, directs it to the workpiece, and provides a vehicle for

directing shielding gas to the weld area. Different types of guns have been designed for many varied applications, ranging from heavy-duty guns for high-current, high-volume production to lightweight guns for low-current or out-of-position welding. The most commonly used guns are designed to be cooled by the surrounding air (Fig. 4). However, as amperage requirements increase, a water-cooled gun may be required. Guns are rated based on their current-carrying capacity, generally with a CO2 shielding gas. If inert gases are used, these gun ratings must be reduced significantly. Guns can also be equipped with their own integral electrode feed units.

FIG. 4 CROSS SECTION OF AN AIR-COOLED GMAW GUN

The contact tube, usually made of copper or a copper alloy, is used to transmit welding current to the electrode, as well as to direct the electrode toward the work. The contact tube is connected electrically to the welding power supply by the power cable. The inner surface of the contact tube is very important, because the electrode must feed easily through this tube while making a good electrical contact. Generally, the hole in the contact tube would be from 0.13 to 0.25 mm (0.005 to 0.010 in.) larger than the wire being used, although larger sizes may be required for materials such as aluminum. The hole should be checked periodically and replaced if it has become elongated because of excessive wear. If a tip in this condition is used, it can result in poor electrical contact and erratic arc characteristics. The nozzle directs an even-flowing column of shielding gas into the welding zone. It is extremely important to maintain an even flow in order to adequately protect the molten weld metal from atmospheric contamination. Different-sized nozzles are available and should be chosen according to the application, that is, larger nozzles for high-current work where the puddle is large and smaller nozzles for low-current and short-circuiting welding. The electrode conduit and liner are connected to a bracket adjacent to the feed rolls on the electrode feed motor. The conduit and liner support both protect and direct the electrode from the feed rolls to the gun and contact tube. Uninterrupted electrode feeding is necessary to ensure good arc stability. A steel liner is recommended when using hard

electrode materials, such as steel and copper, whereas nylon liners should be used for soft electrode materials, such as aluminum and magnesium. The electrode feed unit, or wire feeder, consists of an electric motor, output shaft, drive rolls, and accessories for

maintaining electrode alignment and pressure (Fig. 5). These units can be separate or integrated with the speed control or located remotely from it. The electrode feed motor is usually a direct-current type and provides the mechanical energy for pushing the electrode through the gun and to the work. It has a control circuit that varies the motor speed over a broad range. The feed unit can be an integral component of the gun (Fig. 6) or dual-feed units, one in the gun and one in a separate feeder, can be electrically coupled together to provide a "push-pull" system. The spool-gun and push-pull systems are often used on aluminum, where difficulty can be encountered in trying to push the wire through a conduit to the gun.

FIG. 5 TYPICAL REMOTE-WIRE FEED UNITS. (A) SCHEMATIC OF A TYPICAL FOUR-ROLL SYSTEM MECHANISM. (B) CLOSE-UP VIEW OF A TYPICAL TWO-ROLL SYSTEM GEARBOX

FIG. 6 SPOOL-ON-GUN TYPE TORCH WITH COVER REMOVED TO SHOW SELF-CONTAINED PULL-TYPE FEEDUNIT SYSTEM

The feed motor is connected through a gear reducer to a set of wire feed rolls that transmit mechanical energy to the electrode, pulling it from the source and pushing it through the welding gun. Various types of rolls are available, including knurled, "U" groove, "V" groove, and flat. The knurled design is used for harder wires, such as steel, and allows maximum frictional force to be transmitted to the wire with a minimum of drive roll pressure. These types of rolls are not recommended for soft wires, such as aluminum, because they tend to cause the wire to flake, which can eventually clog the gun or liner. For these softer wires, the "U" groove or "V" groove type will allow the application of uniform pressure around the wire without deforming it. The flat rolls can be used with the smaller-diameter wires and in combination with a "U" or "V" groove. The welding control mechanism and the electrode feed motor for semiautomatic operation are usually provided in one

integrated package. The main function of the welding control mechanism is to regulate the speed of the electrode feed motor, usually through an electronic governor. The control also regulates the starting and stopping of the electrode feed through a signal received from the gun switch.

Normally, shielding gas, water (when used), and welding power are also delivered to the gun through the control mechanism, which requires direct connection to these facilities and the power supply. Gas and water flow are regulated to coincide with the weld start and stop by using solenoid valves. The control mechanism can also sequence gas flow starts and stops and energize the power source contractor. The control mechanism may allow some gas to flow before welding starts (preflow) and after welding stops (postflow) to protect the molten weld puddle. The control mechanism is usually independently powered by 115 V ac. The welding power source provides suitable electrical power (generally 20 to 80 V) that is delivered to the electrode and workpiece to produce the arc. Because the vast majority of GMAW applications utilize DCEP, the positive lead is connected to the gun and the negative lead, to the workpiece. The power source can be the "static" type in which incoming utility power (120 to 480 V) is reduced to welding voltage by a transformer or solid-state inverter. It could also be the "rotating" type in which the welding power is provided by a rotating generator driven by a motor or internal combustion engine. The static type is normally used in shops where there is an available source of power. It has advantages over the rotating type in that it can respond more rapidly to varying arc conditions. The rotating type is generally used at field sites where external power in unavailable. Both types of power sources can be designed and built to provide either a constant current or constant potential (cp) output, the latter of which is the most common by far. On newer power sources, this cp output can be pulsed at either a constant or variable frequency. With the advent of solidstate electronic power sources, such as inverters, even further control over the pulsing variables (for example, frequency, pulse width, and so on) can be obtained.

When used in conjunction with a constant-speed wire feeder, the constant-voltage power source compensates for the variations in the contact-tip-work-distance that can occur during normal welding operations. It does this by instantaneously increasing or decreasing welding current to increase or decrease the electrode burnoff rate. The initial arc length is established by adjusting the voltage at the power source. Once this is set, no other changes are required during welding. The wire feed speed, which is also the current control, is then set by the operator and adjusted as necessary. In addition to this self-regulating feature of the CP power source, control over slope and inductance is included on those machines intended for short-circuiting transfer. Additional controls are also provided when using power sources that have pulsing capabilities. Additional information is available in the article "Power Sources" in this Volume. Electrode Source. The GMAW process uses a continuously fed electrode that is consumed at relatively high speeds.

Therefore, the electrode source must provide a large volume of material that can readily be fed to the gun to ensure maximum process efficiency. This source is usually in the form of a spool or coil that can hold from 7 to 27 kg (15 to 60 lb) of wire that has been wound to allow free feeding without kinks or tangles. Larger spools of up to 115 kg (250 lb) are also available, and material can be provided in drums of 340 to 455 kg (750 to 1000 lb). Small spools of 0.45 to 0.9 kg (1 to 2 lb) are used for spool-on-gun equipment. Regulated Shielding Gas Supply. A system is required to provide constant shielding gas pressure and flow rate during

welding. This system consists of a regulator connected to a supply of "welding grade" shielding gas, as well as the necessary hoses or piping. The regulator is a device that reduces the source gas pressure to a constant working pressure, regardless of variations at the source. It can be a single-stage or dual-stage type and may have a built-in flowmeter. The shielding gas source can be a high-pressure cylinder, a liquid-filled cylinder, or a bulk-liquid tank. Gas mixtures are available in a single cylinder. Mixing devices can also be used to obtain the correct proportions when two or more gases or liquids are used. The type and size of the gas storage source depend on economic considerations that are based on the volume of shielding gas consumed per unit of time. Gas-Metal Arc Welding D.B. Holliday, Westinghouse Electric Corporation

Consumables The two consumable, but essential, elements of the GMAW process are the electrode and the shielding gas, each of which is described below. The chemical composition of the electrode must be selected to achieve the desired properties in the weld metal. The composition is designed with extra deoxidizers or other scavenging agents to compensate for reactions with the

atmosphere and the base metal. The deoxidizers most commonly used in steel electrodes are silicon and manganese. Silicon can also be added in all transfer modes to increase weld metal fluidity or it can be added when a 300 series stainless steel electrode is used. The physical characteristics (finish, straightness, and others) of electrodes used in the GMAW process are important

to successful welding. The material specifications for these electrodes establish manufacturing requirements to ensure that users receive a uniform product that feeds smoothly through the equipment and has these characteristics, as well: • • • •

UNIFORM WINDING ON THE SPOOL OR COIL WITH NO KINKS OR BENDS SMOOTH SURFACE FINISH FREE OF SLIVERS, SCRATCHES, OR SCALE PRESCRIBED CAST AND HELIX UNIFORM DIAMETER

Cast and helix refer to dimensions of a single coil of wire removed from a spool or coil and layed (that is, cast) on a flat surface. If this coil is too small in diameter (cast) or shows an excessive lift from the flat surface (helix) wire, feeding problems during welding can be anticipated. Shielding Gas The primary function of the shielding gas in most of the welding processes is to protect the surrounding atmosphere from contact with molten metal. In the GMAW process, this gas plays an additional role in that it has a pronounced effect on arc characteristics, mode of metal transfer, depth of fusion, weld bead profile, welding speed, and cleaning action. Inert gases, such as argon and helium, are commonly used, as is the active gas, CO2. It is also common to use mixtures of these gases and to employ small additions of oxygen. Information about shielding gas compositions and about which gases to use for specific joining applications is provided in the article "Shielding Gases" in this Volume. Gas-Metal Arc Welding D.B. Holliday, Westinghouse Electric Corporation

Safety The major hazards of concern during GMAW are: the fumes and gases, which can harm health; the high-voltage electricity, which can injure and kill; the arc rays, which can injure eyes and burn skin; and the noise which may be present that can damage hearing. The type and amount of fumes and gas present during welding depend on the electrode being used, the alloy being welded, and the presence of any coatings on the base metal. To guard against potential hazards, a welder should keep his head out of the fume plume and avoid breathing the fumes and gases caused by the arc. Ventilation is always required. Electrode shock can result from exposure to the high open-circuit voltages associated with welding power supplies. All electrical equipment and the workpiece must be connected to an approved electrical ground. Cables should be of sufficient size to carry the maximum current required. Insulation should be protected from cuts and abrasion, and the cable should not come into contact with oils, paints, or other fluids which may cause deterioration. Work areas, equipment, and clothing must be kept dry at all times. The welder should be well insulated, wearing dry gloves, rubbersoled shoes, and standing on a dry board or platform while welding. Radiant energy, especially in the ultraviolet range, is intense during GMAW. To protect the eyes from injury, the proper filter shade for the welding-current level selected should be used. These greater intensities of ultraviolet radiation can cause rapid disintegration of cotton clothing. Leather, wool, and aluminum-coated cloth will better withstand exposure to arc radiation and better protect exposed skin surfaces.

When noise has been determined to be excessive in the work area, ear protection should be used. This can also be used to prevent spatter from entering the ear. Conventional fire prevention requirements, such as removal of combustibles from the work area, should be followed. Sparks, slag, and spatter can travel long distances, so care must be taken to minimize the start of a fire at locations removed from the welding operation. For further information, see the guidelines set forth in the National Fire Protection Association Standard NFPA No. 51B, "Fire Protection in Use of Cutting and Welding Processes." Care should be exercised in the handling, storage, and use of cylinders containing high-pressure and liquefied gases. Cylinders should be secured by chains or straps during handling or use. Approved pressure-reducing regulators should be used to provide a constant, controllable working pressure for the equipment in use. Lubricants or pipe fitting compounds should not be used for making any connections, as they can interfere with the regulating equipment, and in the case of oxygen service, they can contribute to a catastrophic fire. For further safety information, see ANSI/AWS Z49.1, "Safety in Welding and Cutting." Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Introduction IN THE FLUX-CORED ARC WELDING (FCAW) process, the heat for welding is produced by an electric arc between a continuous filler metal electrode and the workpiece. A tubular, flux-cored electrode makes this welding process unique. The flux contained within the electrode can make the electrode self-shielding. Alternatively, an external shielding gas may be required. Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Process Features Flux-cored arc welding has two major variations. The gas-shielded FCAW process (Fig. 1) uses an externally supplied gas to assist in shielding the arc from nitrogen and oxygen in the atmosphere. Generally, the core ingredients in gasshielded electrodes are slag formers, deoxidizers, arc stabilizers, and alloying elements.

FIG. 1 GAS-SHIELDED FLUX-CORED ARC WELDING. SOURCE: REF 1

In the self-shielded FCAW process, the core ingredients protect the weld metal from the atmosphere without external shielding. Some self-shielded electrodes provide their own shielding gas through the decomposition of core ingredients. Others rely on slag shielding, where the metal drops being transferred across the arc and the molten weld pool are protected from the atmosphere by a slag covering. Many self-shielded electrodes also contain substantial amounts of deoxidizing and denitrifying ingredients to help achieve sound weld metal. Self-shielded electrodes can also contain arc stabilizers and alloying elements. Advantages. Because it combines the productivity of continuous welding with the benefits of having a flux present, the

FCAW process has several advantages relative to other welding processes. These advantages include: • • • • •

HIGH DEPOSITION RATES, ESPECIALLY FOR OUT-OF-POSITION WELDING LESS OPERATOR SKILL REQUIRED THAN FOR GAS-METAL ARC WELDING (GMAW) SIMPLER AND MORE ADAPTABLE THAN SUBMERGED ARC WELDING (SAW) DEEPER PENETRATION THAN SHIELDED METAL ARC WELDING (SMAW) MORE TOLERANT OF RUST AND MILL SCALE THAN GMAW

Disadvantages of the FCAW process include: • •

SLAG MUST BE REMOVED FROM THE WELD AND DISPOSED OF MORE SMOKE AND FUME ARE PRODUCED IN FCAW THAN IN THE GMAW AND SAW PROCESSES

• •

FUME EXTRACTION IS GENERALLY REQUIRED EQUIPMENT IS MORE COMPLEX AND MUCH LESS PORTABLE THAN SMAW EQUIPMENT

Reference cited in this section

1. WELDING HANDBOOK, 8TH ED., VOL 2, AWS, 1991 Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Applications Flux-cored arc welding enjoys widespread use in many industries. Both the gas-shielded and self-shielded FCAW processes are used to fabricate structures from carbon and low-alloy steels. Both process variants are used for shop fabrication, but the self-shielded FCAW process is preferred for field use. The acceptability of the FCAW process for structural use is illustrated by the fact that prequalified joints are included in the structural welding code of the American Welding Society (AWS). Gas-shielded flux-cored electrodes are commonly used to weld carbon, low-alloy steel, and stainless steels in the construction of pressure vessels and piping for the chemical processing, petroleum refining, and power-generation industries. In addition, flux-cored electrodes are used to weld some nickel-base alloys. Flux-cored electrodes are also used in the automotive and heavy-equipment industries in the fabrication of frame members, axle housings, wheel rims, suspension components, and other parts. Small-diameter flux-cored electrodes are used for automotive body repair. Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Equipment The FCAW process utilizes semiautomatic, mechanized, and fully automatic welding systems. The basic equipment includes a power supply, wire feed system, and welding gun. The required auxiliary equipment, such as shielding gas, depends on the process variant used and the degree of automation. Fume removal equipment must also be considered in most applications of the FCAW process. Typical semiautomatic equipment is shown in Fig. 2. The equipment used in the gas-shielded FCAW process is typically identical to GMAW equipment.

FIG. 2 SEMIAUTOMATIC FCAW EQUIPMENT. SOURCE: REF 1

The recommended power supply for the semi-automatic FCAW process is a constant-voltage direct current (dc) machine. Most power supplies used for semiautomatic FCAW have output ratings of 600 A or less. A power supply rated at 60% or more duty cycle is the best choice for most industrial applications, whereas a duty-cycle rating as low as 20% may be sufficient for maintenance and repair applications. Constant-current power supplies are used in certain situations, such as field welding applications, where portable constant-current SMAW power supplies are readily available. The addition of a contactor and a voltage-sensing wire feeder makes this an adequate welding system. However, such a system is only recommended when the use of a constantvoltage system is not feasible, because constant-current systems produce an inherently less-stable welding arc than constant-voltage systems. Wire feeders for constant-voltage FCAW systems are generally simple devices that provide a constant wire feed speed. The power supply provides sufficient current to maintain an arc at the voltage that is preset at the power supply. A change in wire feed speed results in a change in the welding current. In a constant-current system, the wire feeder is somewhat more complex. The welding current is preset at the power supply. The wire feeder has a voltage-sensing feedback loop that allows it to adjust the wire feed speed to maintain the desired welding voltage. The wire feeder generally contains systems to close the contactor and open the shielding gas solenoid valve (gas-shielded FCAW process only) when welding is started. Because flux-cored wires are easily deformed by excessive feed roll pressure, knurled feed rolls are generally used in the FCAW process. Some wire feeders use a single drive roll paired with an undriven pressure roll. Others have one or two

pairs of drive rolls. It is generally believed that systems having two pairs of drive rolls require the least drive roll pressure to provide dependable feeding. Both air-cooled and water-cooled welding guns are used in the semiautomatic FCAW process. Air-cooled guns are generally preferred, because they are simpler to maintain, lighter in weight, and less bulky. Water-cooled guns may be required when welding currents over 500 A are used, especially when the shielding gas is rich in argon. Air-cooled guns designed for gas-shielded welding should not be used for self-shielded welding, because the gun depends on the flow of shielding gas for proper cooling. Although curved-neck guns are the most common, straight guns are used to a limited extent. A trigger switch on the welding gun is closed to initiate wire feeding, welding current flow, and shielding gas flow. The electrode is delivered from the wire feeder to the gun through a flexible conduit. Standard conduit lengths are 3, 3.7, 4.6, and 6 m (10, 12, 15, and 20 ft). Other lengths may also be available. Mechanized and automatic FCAW equipment is not substantially different from that used in the semiautomatic

FCAW process. The power supply should be rated for 100% duty cycle. Power supplies capable of outputs up to 1000 A may be required for some applications. Constant-current systems are very seldom used for mechanized and automatic welding. The wire feed system is separated into a drive motor assembly and a welding control device, the latter of which often has a system to automatically start the travel mechanism when wire feed, current flow, and shielding gas flow are initiated. The welding control device is often equipped with a voltmeter and ammeter, as well. The welding gun in mechanized and automatic systems is often mounted directly to the drive motor assembly, eliminating the need for a conduit. The gun is usually straight, but curved-neck guns are also used. Both water- and air-cooled guns are used, depending on the welding current level and shielding gas. Most of the commonly available air-cooled guns can be used at levels up to 500 A with CO2 shielding gas. When argonrich shielding gas is used, the same gun may only be suitable for use at levels up to 300 A. Water-cooled guns are generally used at higher current levels. Various travel mechanisms are used, depending on the applications. These mechanisms include side-beam carriages, tractor-type carriages, and robots. Fume-Removal Equipment. In many cases, the amount of fumes generated by the FCAW process is sufficient to

require fume-removal equipment. Although such equipment can be as simple as exhaust fans in the shop roof, local fume collection is often necessary. These systems can be either collection hoods (located above the welding gun) or fumeextractor guns. These guns are more efficient at collecting fumes, but are heavier and more bulky than standard welding guns. Fume collection hoods must be repositioned each time the welding location is moved in order to be effective.

Reference cited in this section

1. WELDING HANDBOOK, 8TH ED., VOL 2, AWS, 1991 Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Base Metals Most weldable carbon and low-alloy steels can be welded by the FCAW process if suitable electrodes are available. Steels for which electrodes are available include: •

MILD STEELS, SUCH AS AISI 1010 TO 1030, ASTM A 36, A 285, A 515, AND A 516

• • • • • •

WEATHERING STRUCTURAL STEELS, SUCH AS ASTM A 588 HIGH-STRENGTH, LOW-ALLOY STEELS, SUCH AS ASTM A 710 HIGH-TEMPERATURE CHROMIUM-MOLYBDENUM STEELS, SUCH AS ASTM A 387 GRADES 12 (1CR-0.5MO) AND 22 (2.25CR-1MO) NICKEL-BASE STEELS, SUCH AS ASTM A 203 HIGH-STRENGTH QUENCHED AND TEMPERED STEELS, SUCH AS HY-80 AND ASTM A 514 AND A 517 MEDIUM-CARBON, HEAT-TREATABLE, LOW-ALLOY STEELS, SUCH AS AISI 4130

Several grades of stainless steel are welded with both gas-shielded and self-shielded flux-cored electrodes. Electrodes designed for welding AISI types 304, 316, 347, and others are available. Also available are electrodes suitable for joining stainless steels to carbon and low-alloy steels. Electrodes that can be used to weld some nickel-base alloys have also been introduced. Among the nickel-base alloys that have been joined using the FCAW process are alloy 600 (UNS N06600) and alloy 625 (UNS N06625). Some cast irons are also welded using nickel-base flux-cored electrodes designed specifically for this purpose. Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Electrode Manufacture Flux-cored electrodes are generally manufactured using the process shown in Fig. 3. A flat sheath material is first formed into a "U" shape. The core ingredients are poured into this "U" at the desired rate. The sheath is then closed around the core materials to form a round tube. The diameter of this tube is then reduced, generally by drawing or rolling operations, to compress the core materials and bring the electrodes to a size that is usable for welding. The finished wire is than wound on spools, coils, or other packages. Other manufacturing methods, which are generally considered proprietary, are also used.

FIG. 3 FLUX-CORED ELECTRODE MANUFACTURING PROCESS. SOURCE: REF 2

The usability of a flux-cored electrode, as well as the properties and chemistry of the weld it deposits, primarily depends on the ingredients in the electrode core. Common core ingredients and their functions are listed in Table 1.

TABLE 1 FUNCTIONS OF COMMON CORE INGREDIENTS IN FCAW ELECTRODES CORE INGREDIENT

RUTILE (TIO2) FLUORSPAR (CAF2) LIME (CACO3) FELDSPAR SYNTHETIC FRITS MANGANESE SILICON TITANIUM ALUMINUM CHROMIUM, NICKEL, MOLYBDENUM

GAS FORMER

DEOXIDIZER

DENITRIFIER

X

X X X X

SLAG FORMER

VISCOSITY CONTROL

X X X X X

X X

X

ARC STABILIZER

ALLOY

X X X X X

X X X

Reference cited in this section

2. C.E. JACKSON, "FLUXES AND SLAGS IN WELDING," WRC BULLETIN 190, WELDING RESEARCH COUNCIL Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

Electrode Diameters Flux-cored electrodes are produced in diameters ranging from 0.8 to 3.2 mm (0.030 to

1 in.). Electrodes for all-position 8

welding may be available in 0.8 mm (0.030 in.), 0.9 mm (0.035 in.), 1.2 mm (0.045 in.), 1.4 mm (0.052 in.), and 1.6 mm 1 1 5 in.) diameters. Electrodes for flat- and horizontal-position welding may be available in 1.6 mm ( in.), 2.0 mm ( 16 16 64 3 7 1 in.), 2.4 mm ( in.), 2.8 mm ( in.), and 3.2 mm ( in.) diameters. Other diameters may be made available by 32 64 8

(

agreement between the manufacturer and the purchaser. Electrode Classification Carbon and Low-Alloy Steel Electrodes. Carbon steel flux-cored electrodes are classified by AWS A5.20, "Specification for Carbon Steel Electrodes for Flux Cored Arc Welding." The layout of this classification system is defined in Fig. 4; an "X" indicates a position where a designator would be. For example, the designator indicating minimum tensile strength can be a 6, to denote 62 ksi (430 MPa), or a 7, to denote 72 ksi (495 MPa). The assignment of the final designator depends on the shielding requirements and features of the electrode, as well as the recommended welding polarity. These are described in Table 2. The designations EXXT-G and EXXT-GS are designed to allow the classification of electrodes that are not covered by any of the other classifications.

TABLE 2 USABILITY TYPE DESIGNATORS FOR FLUX-CORED ELECTRODES TYPE

SHIELDING(A)

SINGLE OR MULTIPASS(B)

T-1

GAS

MULTIPASS SPRAYLIKE

T-2

GAS

SINGLE

T-3

SELF

SINGLE

T-4

SELF

T-5

GAS

MULTIPASS GLOBULAR 27 J AT -30 °C DCEP(D) (20 FT · LBF AT -20 °F)

T-6

SELF

MULTIPASS SPRAYLIKE

T-7

SELF

MULTIPASS GLOBULAR NO REQUIREMENT

T-8

SELF

MULTIPASS GLOBULAR 27 J AT -30 °C DCEN (20 FT · LBF AT -20 °F)

T-10 SELF

TRANSFER

SPRAYLIKE

IMPACT TOUGHNESS REQUIREMENT

27 J AT -18 °C DCEP (20 FT · LBF AT 0 °F) NO DCEP REQUIREMENT

SPRAYNO LIKE REQUIREMENT MULTIPASS GLOBULAR NO REQUIREMENT

SINGLE

POLARITY(C)

DCEP DCEP

27 J AT -30 °C DCEP (20 FT · LBF AT -20 °F)

GLOBULAR NO REQUIREMENT T-11 SELF MULTIPASS SPRAYNO LIKE REQUIREMENT T-G NOT MULTIPASS NOT NO SPECIFIED SPECIFIED REQUIREMENT TNOT SINGLE NOT NO GS SPECIFIED SPECIFIED REQUIREMENT

DCEN

DCEN DCEN

SPECIAL CHARACTERISTICS

LOW SPATTER, FULL SLAG COVERAGE LOW SPATTER, FULL SLAG COVERAGE, HIGH DEOXIDIZERS HIGH SPEED HIGH DEPOSITION, LOW PENETRATION, CRACK RESISTANT IMPROVED TOUGHNESS, CRACK RESISTANT, THIN SLAG IMPROVED TOUGHNESS, DEEP PENETRATION CRACK RESISTANT, GOOD SLAG REMOVAL IMPROVED TOUGHNESS, CRACK RESISTANT HIGH SPEED GENERAL PURPOSE NOT SPECIFIED

NOT SPECIFIED NOT NOT SPECIFIED SPECIFIED

Source: AWS A5.20

(A) GAS-SHIELDED ELECTRODES ARE CLASSIFIED USING CO2 SHIELDING; HOWEVER, AR-CO2 MIXTURES ARE COMMONLY USED IN PRACTICE. (B) MULTIPASS ELECTRODES ARE SUITABLE FOR SINGLE-PASS OR MULTIPASS WELDING; SINGLE-PASS ELECTRODES ARE SUITABLE FOR SINGLE-PASS WELDING ONLY.

(C) DCEP, DIRECT-CURRENT ELECTRODE POSITIVE; DCEN, DIRECT-CURRENT ELECTRODE NEGATIVE. (D) T-5 ELECTRODES ARE CLASSIFIED USING DCEP; HOWEVER, THEY ARE SOMETIMES USED WITH DCEN IN PRACTICE.

FIG. 4 CLASSIFICATION SYSTEM FOR CARBON STEEL FLUX-CORED ELECTRODES. THE LETTER "X" AS USED IN THIS FIGURE AND IN ELECTRODE CLASSIFICATION DESIGNATIONS IN AWS SPECIFICATION A5.20-79 SUBSTITUTES FOR SPECIFIC DESIGNATIONS INDICATED BY THIS FIGURE. SOURCE: REF 3

It should be noted that the Canadian Standards Association specification W48.5-M1990, "Carbon Steel Electrodes for Flux- and Metal-Cored Arc Welding," includes two additional electrode types that are planned for inclusion in the next revision of the AWS A5.20 specification. These classifications are EXXT-9 and EXXT-12. T-9 electrodes are generally similar to T-1 electrodes, but with improved toughness. T-12 electrodes are generally similar to T-9 electrodes, but with a maximum tensile strength also specified. Low-alloy steel electrodes are classified under AWS A5.29, "Specification for Low Alloy Steel Electrodes for Flux Cored Arc Welding." This specification uses a classification system very similar to that used in specification AWS A5.20, except that a chemical composition designator is added to the designation. An electrode classified according to this specification will have the form EXXTX-X. All of the positions before the dash have the same meaning as those in specification AWS A5.20. The position behind the dash is the chemical composition designator, which consists of a letter and a number. The letter denotes the alloy type of the electrode as follows:

DESIGNATOR A B NI D W K G

ALLOY TYPE CARBON-MOLYBDENUM STEEL CHROMIUM-MOLYBDENUM STEEL NICKEL STEEL MANGANESE-MOLYBDENUM STEEL WEATHERING STEEL OTHER LOW-ALLOY STEELS NOT SPECIFIED

Specification AWS A5.29 classifies only the EXXT1-X, EXXT4-X, EXXT5-X, and EXXT8-X electrode types, and the usability designator has the name meaning as that in specification A5.20. Minimum tensile strengths of up to 830 MPa (120 ksi) are included in specification AWS A5.29. Impact toughness requirements are based on the strength, usability, and chemical composition requirements of the electrode.

Stainless steel electrodes are classified under AWS A5.22, "Specification for Flux Cored Corrosion Resisting

Chromium and Chromium-Nickel Steel Electrodes." Classifications to this specification have the form EXXXT-X. The first three positions are the chemical composition designator, which corresponds to the American Iron and Steel Institute (AISI) designations (such as 308, 316, and 410) of steels having a similar composition. The final position is the shieldingtype designator. T-1 types are designed for use with CO2 or Ar-CO2 shielding gases (classification requires the use of CO2). T-2 types are designed for use with Ar-2O2 shielding gas. T-3 types are self-shielded. A T-G type is included for electrodes that are not covered by the other shielding-type designators. Nickel-Base Electrodes. A specification for nickel-base flux-cored electrodes has not yet been published. The

manufacturers of these electrodes should be consulted for information on their usability, composition, and properties.

Reference cited in this section

3. "SPECIFICATION FOR CARBON STEEL ELECTRODES FOR FLUX CORED ARC WELDING," AWS A5.20-79, AWS, 1979 Flux-Cored Arc Welding David W. Meyer, The Esab Group, Inc.

References

1. WELDING HANDBOOK, 8TH ED., VOL 2, AWS, 1991 2. C.E. JACKSON, "FLUXES AND SLAGS IN WELDING," WRC BULLETIN 190, WELDING RESEARCH COUNCIL 3. "SPECIFICATION FOR CARBON STEEL ELECTRODES FOR FLUX CORED ARC WELDING," AWS A5.20-79, AWS, 1979 Gas-Tungsten Arc Welding Grant Ken-Hicken, Sandia National Laboratory

Introduction GAS-TUNGSTEN ARC WELDING (GTAW), also known as HeliArc, tungsten inert gas (TIG), and tungsten arc welding, was developed in the late 1930s when a need to weld magnesium became apparent. Russell Meredith (Ref 1) developed a welding process using the inert gas helium and a tungsten electrode to fuse magnesium. This joining method replaced riveting as a method of building aircraft with aluminum and magnesium components. The HeliArc welding has continued to this day with many refinements and name changes, but with no change in the fundamentals demonstrated by Meredith (Ref 1). The melting temperature necessary to weld materials in the GTAW process is obtained by maintaining an arc between a tungsten alloy electrode and the workpiece (Fig. 1). Weld pool temperatures can approach 2500 °C (4530 °F). An inert gas sustains the arc and protects the molten metal from atmospheric contamination. The inert gas is normally argon, helium, or a mixture of helium and argon.

FIG. 1 SCHEMATIC SHOWING KEY COMPONENTS AND PARAMETERS OF THE GTAW PROCESS. SOURCE: REF 2

References

1. R. MEREDITH, U.S. PATENT 2,274,631 2. S. KOU, WELDING METALLURGY, JOHN WILEY & SONS, 1987, P 10 Gas-Tungsten Arc Welding Grant Ken-Hicken, Sandia National Laboratory

Applications Gas-tungsten arc welding is used extensively for welding stainless steel, aluminum, magnesium, copper, and reactive materials (for example, titanium and tantalum). The process can also be used to join carbon and alloy steels. In carbon steels, it is primarily used for root-pass welding with the application of consumable inserts or open-root techniques on pipe. The materials welded range from a few thousandths of an inch to several inches in thickness. Advantages and Limitations Advantages of GTAW include (Ref 3): • • • • • •

PRODUCES HIGH-QUALITY, LOW-DISTORTION WELDS FREE OF THE SPATTER ASSOCIATED WITH OTHER METHODS CAN BE USED WITH OR WITHOUT FILLER WIRE CAN BE USED WITH A RANGE OF POWER SUPPLIES WELDS ALMOST ALL METALS, INCLUDING DISSIMILAR ONES GIVES PRECISE CONTROL OF WELDING HEAT

The GTAW process is applicable when the highest weld quality is required. It can be used to weld almost all types of metals. The operator has excellent control of heat input, and vision is not limited by fumes or smoke from the process. Limitations of GTAW include (Ref 4):

• •

•

PRODUCES LOWER DEPOSITION RATES THAN CONSUMABLE ELECTRODE ARC WELDING PROCESSES REQUIRES SLIGHTLY MORE DEXTERITY AND WELDER COORDINATION THAN GAS METAL ARC WELDING (GMAW) OR SHIELDED METAL ARC WELDING (SMAW) FOR MANUAL WELDING LESS ECONOMICAL THAN CONSUMABLE ELECTRODE ARC WELDING FOR THICK SECTIONS GREATER THAN 9.5 MM (

•

3 8

IN.)

PROBLEMATIC IN DRAFTY ENVIRONMENTS BECAUSE OF DIFFICULTY IN SHIELDING THE WELD ZONE PROPERLY

Additional problems with the process may include: • • • • •

TUNGSTEN INCLUSIONS IF THE ELECTRODE IS ALLOWED TO CONTACT THE WELD POOL CONTAMINATION OF THE WELD METAL, IF PROPER SHIELDING OF THE FILLER METAL BY THE GAS STREAM IS NOT MAINTAINED LOW TOLERANCE FOR CONTAMINANTS ON FILLER OR BASE METALS CONTAMINATION OR POROSITY, CAUSED BY COOLANT LEAKAGE FROM WATERCOOLED TORCHES ARC BLOW OR ARC DEFLECTION, AS WITH OTHER PROCESSES

Power supplies for GTAW are usually the constant-current type with a drooping (negative) volt-ampere (V-A) curve.

Saturable reactors and thyristor-controlled units are the most common. Advances in the electronics industry have readily been accepted in the welding community, resulting in sophisticated, lightweight power supplies. Transistorized direct current (dc) power supplies are becoming common, and the newer rectifier-inverter supplies are very compact and versatile. The inverter power supply consists of three converters:

• • •

60 HZ PRIMARY ALTERNATING CURRENT (AC) IS RECTIFIED TO DC DIRECT CURRENT IS INVERTED TO HIGH-FREQUENCY AC ALTERNATING CURRENT IS RECTIFIED TO DC (REF 5)

The inverter supplies can be switched from constant current to constant voltage for GMAW, resulting in a very versatile piece of equipment. The inverter-controlled power supplies are more stable and have faster response times than conventional silicon-controlled rectifier (SCR) power supplies. Figure 2 compares the response of an inverter-controlled arc welding machine and a thyristor-controlled welding machine.

FIG. 2 STARTING CURRENT WAVEFORMS OF TWO POWER SOURCES TO SHOW RELATIVE RESPONSE TIMES OF EACH SOURCE. (A) THYRISTOR-CONTROLLED SOURCE. (B) INVERTER-CONTROLLED SOURCE. FASTER RESPONSE OF INVERTER-CONTROLLED ARC WELDING MACHINE (2 MS TO GO FROM 0 TO 100 A) INDICATES A MORE STABLE ARC.

Torch Construction. The welding torch holds the tungsten electrode that conducts the current to the arc, and it provides a means of shielding the arc and molten metal. The major components of a typical welding torch are shown in Fig. 3.

FIG. 3 SCHEMATIC SHOWING EXPLODED VIEW OF KEY COMPONENTS COMPRISING A GTAW MANUAL TORCH

Welding torches rated at less than 200 A are normally gas-cooled (that is, the shielding gas flows around the conductor cable, providing the necessary cooling). Water-cooled torches are used for continuous operation or at higher welding currents and are common for mechanized or automatic welding (see the section "GTAW Process Variations" in this article). The cooling water may be supplied to the torch from a recirculating tank that uses a radiator or chiller to cool the water. Electrodes. The nonconsumable electrodes used in GTAW are composed of tungsten or alloys of tungsten. The most

common electrode is a 2% ThO2-W alloy (EWTh-2). This material has excellent operating characteristics and good stability. Thoria is radioactive, so care must be taken when sharpening electrodes not to inhale metal dust. The grindings are considered hazardous waste in some states, and disposal may be subject to environmental regulations. Lanthaniated (EWLa-1) and yttriated tungsten electrodes have the best starting characteristics in that an arc can be started and maintained at a lower voltage. Ceriated tungsten (EWCe-2) is only slightly better than the thoriated tungsten with respect to arc starting and melt-off rate. Any of the aforementioned electrodes produce acceptable welds. The easy starting of the lanthaniated electrode is a result of the lower work function (Ref 6) which allows it to emit electrodes readily at a lower voltage. Pure tungsten is used primarily in ac welding and has the highest consumption rate. Alloys of zirconium are also used. Tungsten electrodes are classified on the basis of their chemical composition (Table 1). Requirements for tungsten

electrodes are given in the latest edition of ANSI/AWS A5.12 ("Specification for Tungsten and Tungsten Alloy Electrodes for Arc Welding and Cutting"). The shape of the electrode tip can affect the resulting weld shape. Electrodes with included angles from 60 to 120° are stable and give good weld penetration depth-to-width ratios (Ref 7). Electrodes with smaller included angles (5 to 30°) are used for grooved weld joints to eliminate arcing to the part side walls.

TABLE 1 CLASSIFICATION OF ALLOYING ELEMENTS IN SELECTED TUNGSTEN ALLOY ELECTRODES FOR GTAW APPLICATIONS AWS CLASSIFICATION

COLOR(A)

ALLOYING ELEMENT

ALLOYING OXIDE

EWP EWCE-2 EWLA-1 EWTH-1 EWTH-2 EWZR-1 EWG

GREEN ORANGE BLACK YELLOW RED BROWN GRAY

... CERIUM LANTHANUM THORIUM THORIUM ZIRCONIUM NOT SPECIFIED(B)

... CEO2 LA2O3 THO2 THO2 ZRO2 ...

ALLOYING OXIDE, WT%

2 1 1 2 0.25 ...

Source: Ref 4

(A) COLOR MAY BE APPLIED IN THE FORM OF BANDS, DOTS, AND SO ON, AT ANY POINT ON THE SURFACE OF THE ELECTRODE. (B) MANUFACTURER MUST IDENTIFY THE TYPE AND NOMINAL CONTENT OF THE RAREEARTH OXIDE ADDITION. Wire feed systems are made from a number of components and vary from simple to complex. The basic system consists of a means of gripping the wire sufficiently to pull it from the spool and push it through the guide tube to the point of welding. Electronic switches and controls are necessary for the electric drive motor. The wire will be fed into the leading edge for cold wire feeds and into the trailing edge for hot wire feeds. Cables, hoses, and gas regulators are necessary to deliver the process consumable of electricity, water, and inert gas

to the welding torch. Arc oscillation is used in both manual and mechanized welding (see the section "GTAW Process Variations" in this

article). The benefits in manual welding are basic to the control of the weld when adapting to changes in the weld joint and gap. In mechanized welding, the oscillation is typically produced by moving the entire welding torch mechanically or by moving the arc plasma with the aid of an externally applied magnetic field. Oscillation allows the welding heat to be placed at precise locations. This is advantageous when welding irregularly shaped parts. The number of welding passes and total heat input can be decreased when arc oscillation is used, because it reduces the cost as well as the weld shrinkage and upsetting. Figure 4 shows the effect of magnetic oscillation on distortion. Some alloys need the tempering produced by multipass stringer bead welding techniques, and arc oscillation should not be used. Externally applied magnetic fields can be used to stabilize the arc, minimize arc blow, and displace the arc plasma forward of the welding torch. This results in improved weld appearances and increased welding speeds in tube mills and other high-speed applications (Ref 9).

FIG. 4 EFFECT OF MAGNETIC OSCILLATION ON WELD DISTORTION IN 25 MM (

1 IN.) THICK STAINLESS 2

STEEL. (A) WELD PRODUCED WITH ARC OSCILLATION. (B) STRINGER BEADS PRODUCED IN WELD WITHOUT ARC OSCILLATION. NOTE MINIMAL UPSETTING WITH ARC OSCILLATION. SOURCE: REF 8

References cited in this section

3. T. MYERS, WHY THE GROWING INTEREST IN GAS TUNGSTEN ARC WELDING? THE FABRICATOR, VOL 22 (NO. 9), NOV 1992, P 38 -39 4. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 74-107 5. T. BYRD, INVERTER POWER SOURCES: AN EFFICIENT ALTERNATIVE, WELD. J., VOL 72 (NO. 1), JAN 1993, P 37-40 6. A.A. SADEK, M. USHIO, AND F. MATSUDA, EFFECTS OF RARE EARTH METAL OXIDE ADDITIONS TO TUNGSTEN ELECTRODES, METALL. TRANS. A, VOL 21, 1990, P 3221-3234 7. J.R. KEY, ANODE/CATHODE GEOMETRY AND SHIELDING GAS INTERRELATIONSHIPS IN GTAW, WELD. J., VOL 59 (NO. 12), DEC 1980, P 364-S TO 370-S 8. G.K. HICKEN, N.D. STUCK, AND H.W. RANDAL, APPLICATIONS OF MAGNETICALLY CONTROLLED ARCS, WELD. J., VOL 55 (NO. 4), 1976, P 264-267 9. G.K. HICKEN, C.E. JACKSON, THE EFFECTS OF APPLIED OF MAGNETIC FIELDS ON WELDING ARCS, WELD. J., VOL 45 (NO. 11), NOV 1966, P 515-S TO 524-S Gas-Tungsten Arc Welding Grant Ken-Hicken, Sandia National Laboratory

Process Parameters Welding Current. Current is one of the most important operating conditions to control in any welding operation,

because it is related to the depth of penetration, welding speed, deposition rate, and quality of the weld. Fundamentally, there are but three choices of welding current:

• • •

DIRECT CURRENT ELECTRODE NEGATIVE (DCEN) DIRECT CURRENT ELECTRODE POSITIVE (DCEP) ALTERNATING CURRENT

Figures 5 and 6 show the effect of dc and ac on weld shape. Table 2 gives the recommended current relative to workpiece material.

TABLE 2 SUITABILITY OF TYPES OF CURRENT FOR GTAW OF SELECTED METALS

METAL WELDED

ALTERNATING DCEN DCEP CURRENT(A)

LOW CARBON STEEL: 0.38-0.76 MM (0.015-0.030 IN.)(A) G(B) 0.76-3.18 MM (0.030-0.125 IN.) NR

E E

NR NR

G(B) G(B) G(B) G(B) NR

E E E E E

NR NR NR NR NR

E E E G(B)

NR(C) NR(C) NR(C) E

G NR NR NR

G(B) NR NR

E E E

NR NR NR

E

NR(C)

G

E

NR(C)

NR

E G(B) NR

NR(C) E E

NR NR NR

HIGH-CARBON STEEL CAST IRON STAINLESS STEEL HEAT-RESISTANT ALLOYS REFRACTORY METALS ALUMINUM ALLOYS: 0.64 MM (0.025 IN.) >0.64 MM (0.025 IN.) CASTINGS BERYLLIUM COPPER AND ALLOYS: BRASS DEOXIDIZED COPPER SILICON BRONZE MAGNESIUM ALLOYS: ≤

3.2 MM (

>4.8 MM (

1 8

3 16

IN.) IN.)

CASTINGS SILVER TITANIUM ALLOYS

Note: E, excellent; G, good; NR, not recommended.

(A) STABILIZED. DO NOT USE ALTERNATING CURRENT ON TIGHTLY JIGGED ASSEMBLIES. (B) AMPERAGE SHOULD BE ABOUT 25% HIGHER THAN WHEN DCEN IS USED. (C) UNLESS WORK IS MECHANICALLY OR CHEMICALLY CLEANED IN THE AREAS TO BE WELDED

FIG. 5 EFFECT OF POLARITY ON GTAW WELD CONFIGURATION WHEN USING DIRECT CURRENT. (A) DCEN. DEEP PENETRATION, NARROW MELTED AREA, APPROXIMATE 30% HEAT IN ELECTRODE AND 70% HEAT IN BASE METAL. (B) DCEP. SHALLOW PENETRATION, WIDE MELTED AREA, APPROXIMATE 70% HEAT IN ELECTRODE AND 30% HEAT IN BASE METAL

FIG. 6 WELD CONFIGURATION AS A FUNCTION OF TYPE OF CURRENT (AC OR DC) USED

Alternating current is characterized as reversing the polarity of the work and electrode at 60 Hz. The rapidly changing polarity gives a cathodic cleaning action that is beneficial for oxide removal when welding aluminum and magnesium. The alternating currents result in electrode heating during the DCEP portion of each cycle. This necessitates the use of larger-diameter electrodes, normally made of pure tungsten. Variable polarity welding allows the frequency of polarity switching to be preset. This can produce the cleaning effects to ac welding and the high efficiency of dc welding. Direct current electrode negative is most often used in the GTAW process. This results in maximum application of heat to the work and maximum melting of the workpiece. Pulsed versus Nonpulsed Current. Nonpulsed or continuous current is the standard for GTAW. However, there are several advantages to using pulsed current. Pulsing produces the maximum amount of penetration while minimizing the total heat applied to the part. Pulsing also aids in timing the motion necessary in manual welding and allows the weld pool to cool between pulses. Microwelding refers to a class of weldments that are made at welding currents from 1 to 20 A. In most cases, the welding

is used for electronic applications, bellows, wires, and other components where heat input must be precisely controlled. Shielding Gases. The original GTAW process used helium as the shielding gas for welding magnesium and aluminum.

Today, argon is the predominant shielding gas. Argon is the least expensive of the inert gases used for shielding gas-tungsten arc welds, which is only partially

responsible for its widespread use. Argon has a low ionization potential (2.52 × 10-18 J, or 15.7 eV), making it easier to form an arc plasma than with other shielding gases. Argon is approximately 1.4 times heavier than air, so it displaces air, resulting in excellent shielding of the molten weld pool. Helium has an ionization potential of 3.92 × 10-18 J (24.5 eV), which results in more difficult arc initiation and operation

at a higher arc voltage. The higher arc voltage, V, results in a higher heat input, Q, for a given arc length and current, I:

Q = IVT

(EQ 1)

where Q is in Joules, I is in amperes; and t is in seconds. This high heat input can be very beneficial when welding copper, aluminum, and other high-conductivity materials. Helium shielding used with DCEN is very effective for welding thick aluminum. Gas Purity. Most materials can be welded using a welding grade torch gas with a purity of 99.995% or 50 ppm

impurities. However, some reactive materials (for example, titanium, molybdenum, and tantalum) require that the contaminant level be less than 50 ppm, which may require certified purity or the use of gas filters and purifiers (Ref 10). Gas Flow Rates. Helium, because of its low density, must be used at higher flow rates than argon. Typical flow rates for

argon are 7 L/min (15 ft3/h) and 14 L/min (30 ft3/h) for helium. Backup Purge. Protecting the molten weld pool from the atmosphere is very important in GTAW. Atmospheric

contamination can result in weld cracks, porosity, scaling, and an unacceptable granular appearance. The gas cup on the welding torch is the primary outlet of shielding gas for most GTAW applications. Back side shielding is important because the presence of oxygen can reduce weld metal penetration and result in the effects mentioned above (Ref 11). Balloon and water-soluble paper dams are sometimes used to minimize the volume to be purged. Copper backing bars and ceramics are sometimes used to hold shielding gas against the back surface of the molten weld and support the molten underbead. Reactive materials and special applications may require more elaborate shielding. This can be in the form of a simple trailing device providing the inert shielding gas or may be as elaborate as a special welding chamber equipped with gas purifiers and analyzers. Specially constructed plastic bags have been used successfully to weld large, irregularly shaped components (Fig. 7).

FIG. 7 PLASTIC BAG ENCLOSURE THAT SIMULATES A GLOVE BOX, USED TO PURGE IRREGULARLY SHAPED COMPONENTS IN GTAW OPERATIONS

Filler Metals. The thickness of the part to be welded will determine the need for filler metal additions. Material thinner than 3.2 mm (0.125 in.) can be successfully welded without filler metal additions. Filler metal, when needed, can be added manually in straight length or automatically from a roll or coil. The filler metal is normally added cold; hot wire can be used for automatic applications (Fig. 8). A welding insert is preplaced filler material of several possible configurations to aid in root-pass welding.

FIG. 8 SCHEMATIC SHOWING KEY COMPONENTS AND PARAMETERS OF A GTAW HOT WIRE SYSTEM. SOURCE: REF 4

Rods. Straight lengths of filler wire, typically 915 mm (36 in.) in length, are used for manual welding. Most straight

lengths are round in cross section, but some aluminum fillers are somewhat rectangular. Cold Wire. Coiled wire may be acquired in small 100 mm (4 in.) spools. Larger 305 mm (12 in.) spools or large coils can

weigh over 225 kg (500 lb). The larger coils are normally used with GMAW because it requires larger quantities of filler metal. The filler wire is fed into the leading edge of the weld pool during cold-wire welding. Hot wire GTAW utilizes a heated filler metal to increase the deposition rate of the process (Fig. 9). The wire is resistance heated to near the melting temperature and fed into the trailing edge of the weld pool. Deposition rates to 29 kg/h (65 lb/h) are achievable. The higher deposition rates obtained with hot wire make the process competitive for welds and overlays and improve productivity. Table 3 lists typical hot wire parameters.

TABLE 3 TYPICAL PARAMETERS FOR AUTOMATIC HOT WIRE GTAW ARC CURRENT, A

300 400 500

ARC VOLTAGE, V

TORCH TRAVEL SPEED

WIRE FEED RATE(A)

DEPOSITION RATE

10-12 11-13 12-15

mm/min 100-255 150-355 205-510

mm/min 2790-9400 4700-11300 7490-16900

kg/h 1.4-4.5 2.3-5.4 3.6-8.2

Note: Using a 4.0 to 4.8 mm (

(A)

in./min 4-10 6-14 8-20

in./min 110-370 185-445 292-665

5 3 to in.) diameter 2% Th tungsten electrode with a 75He-25Ar shielding gas. 32 16

WIRE DIAMETER: 1.14 MM (0.045 IN.)

lb/h 3-10 5-12 8-18

FIG. 9 DEPOSITION RATES FOR GTAW WITH COLD AND HOT FILLER WIRE ON A STEEL WORKPIECE

Welding inserts are used to produce a smooth uniform underbead. The insert is normally a separate piece of material, although integral inserts are sometimes used. Smooth underbead can be produced in a mechanized system or in manual systems when highly skilled welders are available. The insert can be obtained in several different configurations (Fig. 10). The purpose of the insert is to preplace the filler metal at the joint root. Inserts with compositions that differ from that of the base metal can be used to improve the weldability of some materials.

FIG. 10 SELECTED JOINT CONFIGURATIONS AND WELDING INSERTS USED IN GTAW PROCESSES. (A) Y-TYPE

5 IN., FOR INDEXING WHEN ACCESSIBILITY IS LIMITED; NO LAND REQUIRED. (B) COMPOUND 32 1 1 1 5 BEVEL WITH OPEN ROOT, IN. GAP; LAND, IN. THICK. WASHER-TYPE INSERT ( × IN.) SHOULD BE 8 16 8 32 1 USED IF INSERT REQUIRED. (C) INSIDE DIAMETER/OUTSIDE DIAMETER PREPARATION WITH OPEN ROOT, 8 3 1 5 IN. GAP; LAND, IN. THICK. WASHER-TYPE INSERT ( × IN.) SHOULD BE USED IF INSERT REQUIRED. 32 8 32 1 5 1 1 (D) WASHER-TYPE INSERT ( × IN.); LAND, IN. THICK. (E) J PREP MUSHROOM INSERT; LAND IN. 8 32 16 16 3 THICK; EXTENDED LAND, IN. LONG 32 INSERT,

References cited in this section

4. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 74-107 10. K.F. KRYSIAK AND P.M. BHADHA, SHIELDING GAS PURIFICATION IMPROVES WELD QUALITY, WELD. J., NOV 1990, P 47-49 11. H. GIEPL, ORBITAL TIG (GTA) WELDING, STAINL. STEEL LEUR., MAY 1992, P 48-55 Gas-Tungsten Arc Welding Grant Ken-Hicken, Sandia National Laboratory

GTAW Process Variations Manual welding refers to the GTAW process in which the welder manipulates the welding torch by hand. If a motorized wire feeder is attached to the torch, the process is classified as semiautomatic welding.

Products generated by skilled manual welders account for a large proportion of GTAW applications. The equipment can be quite inexpensive, and properly trained welders can join a wide variety of materials. Manual welding is used extensively in stainless steel piping as well as for the root pass in carbon steel pipe welds. Mechanized welding may require adjustment to welding parameters in response to visual observation of the weld.

Machine or mechanized welding requires some specialized accessories. The basic system contains a means for holding and moving the welding torch as well as the workpiece. Because arc voltage is an essential variable in GTAW and is proportional to arc length, voltage feedback devices are often used with motorized torch holders to control the arc length. Narrow groove welding makes use of the GTAW cold wire welding process with a narrowed weld joint. Figure 11 shows a typical weld. Narrow groove welding is limited to mechanized welding applications where precise torch location can be maintained.

FIG. 11 TYPICAL NARROW GROOVE WELD PRODUCED IN MECHANIZED WELDING APPLICATIONS

Automatic welding does not require manual 0parameter adjustment or observation of the weld during the welding process. The most common application of automatic welding is associated with orbiting weld heads used to weld pipe and tubing. These devices attach to the workpieces and move around the circumference, fusing the metal. Most systems perform an autogenous weld; others have wire feed and oscillation capabilities. These systems are often used in conjunction with a computer to control the welding variables. Automatic controls utilizing microprocessors and computer numerical control servodrives make it possible to use one welding system to weld a variety of materials and shapes. Figure 12 shows 14 different parts welded with an automatic GTAW system.

WELDING SPECIFICATIONS FOR COMPONENTS SHOWN Welds were made with a 250 A water-cooled GTAW torch that used 1.6 mm (0.062 in.) diameter electrode 6061 ALUMINUM WORKPIECE(A) COMPONENT NO. 1

WORKPIECE TUBE DIAMETER, 22.22 (0.875) MM (IN.) WALL THICKNESS, 1.57 (0.062)

COMPONENT NO. 2

COMPONENT NO. 3

28.58 (1.125) 0.88 (0.035)

26.82 (1.056) 1.65 (0.065)

TYPE 304 STAINLESS STEEL WORKPIECE(B) COMPONENT COMPONENT NO. 4 NO. 5

12.7 (0.5)

19.1 (0.75)

0.63 (0.025)

0.63 (0.025)

MM (IN.) POWER SUPPLY PEAK CURRENT, A BACKGROUND, A PULSE RATE, PPS(C) WIDTH, % TORCH WELD TRAVEL SPEED, MM/MIN WELDING CYCLE TIME, S SHIELDING GAS FLOW RATE, L/MIN (FT3H) FILLER METAL DIAMETER, MM (IN.) FEED RATE, MM/MIN (IN./MIN) JOINT TYPE

50-100 15 5

52-60 15 6

70-85 15 6

20-45 15 8

35-45 15 8

65

65

65

75

75

405 (16)

405 (16)

405 (16)

380 (15)

380 (15)

8.0

4.5

6.5

6.0

10.0

8.5 (18)

8.5 (18)

8.5 (18)

7.1 (15)

7.1 (15)

0.8 (0.03)

0.8 (0.03)

0.8 (0.03)

...

...

3810 (150)

3560 (140)

3810 (140)

...

...

SADDLE

FILLET

FILLET/BUTT

FILLET/BUTT BUTT

(A) 4043 ALUMINUM FILLER METAL; AC/DC SQUARE WAVE POWER SUPPLY; 2% CERIATED ELECTRODE MATERIAL; 50% HE-50% AR SHIELDING GAS; AC PROCESS (B) NO FILLER METAL USED; DC PRECISION POWER SUPPLY; 2% THORIATED ELECTRODE MATERIAL; ARGON SHIELDING GAS; DC PROCESS (C) PULSES PER SECOND FIG. 12 TYPICAL COMPONENTS PRODUCED BY AUTOMATIC GTAW PROCESS. SOURCE: REF 12

Reference cited in this section

12. K.J. PFAHL, AUTOMATIC CONTOUR WELDING OF TUBE AND PIPE, TUBE PIPE Q., VOL 3 (NO. 4), 1992, P 58-62 Gas-Tungsten Arc Welding Grant Ken-Hicken, Sandia National Laboratory

References

1. R. MEREDITH, U.S. PATENT 2,274,631 2. S. KOU, WELDING METALLURGY, JOHN WILEY & SONS, 1987, P 10 3. T. MYERS, WHY THE GROWING INTEREST IN GAS TUNGSTEN ARC WELDING? THE FABRICATOR, VOL 22 (NO. 9), NOV 1992, P 38 -39 4. WELDING HANDBOOK, 8TH ED., VOL 2, AMERICAN WELDING SOCIETY, 1991, P 74-107 5. T. BYRD, INVERTER POWER SOURCES: AN EFFICIENT ALTERNATIVE, WELD. J., VOL 72 (NO. 1), JAN 1993, P 37-40 6. A.A. SADEK, M. USHIO, AND F. MATSUDA, EFFECTS OF RARE EARTH METAL OXIDE ADDITIONS TO TUNGSTEN ELECTRODES, METALL. TRANS. A, VOL 21, 1990, P 3221-3234

7. J.R. KEY, ANODE/CATHODE GEOMETRY AND SHIELDING GAS INTERRELATIONSHIPS IN GTAW, WELD. J., VOL 59 (NO. 12), DEC 1980, P 364-S TO 370-S 8. G.K. HICKEN, N.D. STUCK, AND H.W. RANDAL, APPLICATIONS OF MAGNETICALLY CONTROLLED ARCS, WELD. J., VOL 55 (NO. 4), 1976, P 264-267 9. G.K. HICKEN, C.E. JACKSON, THE EFFECTS OF APPLIED OF MAGNETIC FIELDS ON WELDING ARCS, WELD. J., VOL 45 (NO. 11), NOV 1966, P 515-S TO 524-S 10. K.F. KRYSIAK AND P.M. BHADHA, SHIELDING GAS PURIFICATION IMPROVES WELD QUALITY, WELD. J., NOV 1990, P 47-49 11. H. GIEPL, ORBITAL TIG (GTA) WELDING, STAINL. STEEL LEUR., MAY 1992, P 48-55 12. K.J. PFAHL, AUTOMATIC CONTOUR WELDING OF TUBE AND PIPE, TUBE PIPE Q., VOL 3 (NO. 4), 1992, P 58-62 Plasma Arc Welding Ian D. Harris, Edison Welding Institute

Introduction PLASMA ARC WELDING (PAW) can be defined as a gas-shielded arc welding process where the coalescence of metals is achieved via the heat transferred by an arc that is created between a tungsten electrode and a workpiece. The arc is constricted by a copper alloy nozzle orifice to form a highly collimated arc column (Fig. 1). The plasma is formed through the ionization of a portion of the plasma (orifice) gas. The process can be operated with or without a filler wire addition.

FIG. 1 PLASMA ARC WELDING PROCESS, SHOWING CONSTRICTION OF THE ARC BY A COPPER NOZZLE AND A KEYHOLE THROUGH THE PLATE

Principles of Operation. Once the equipment is set up and the welding sequence is initiated, the plasma and shielding

gases are switched on. A pilot arc is then struck between a tungsten alloy electrode and the copper alloy nozzle within the torch (nontransferred arc mode), usually by applying a high-frequency open-circuit voltage. When the torch is brought in close proximity to the workpiece or when the selected welding current is initiated, the arc is transferred from the electrode to the workpiece through the orifice in the copper alloy nozzle (transferred arc mode), at which point a weld pool is formed (Fig. 1).

The PAW process can be used in two distinct operating modes, often described as the melt-in mode and the keyhole mode. The melt-in-mode refers to a weld pool similar to that which typically forms in the gas-tungsten arc welding (GTAW)

process, where a bowl-shaped portion of the workpiece material that is under the arc is melted. In the keyhole mode, the arc fully penetrates the workpiece material, forming a nominally concentric hole, or keyhole, through the thickness. The molten weld metal flows around the arc and resolidifies behind the keyhole as the torch traverses the workpiece. Current and Operating Modes. The PAW process uses three current modes: microplasma (melt-in mode), medium-

current plasma (melt-in mode), and keyhole plasma (keyhole mode). This categorization is primarily based on the level of welding current. The microplasma mode is usually defined in the current range from 0.1 to 15 A. The medium-current plasma mode ranges from 15 to 100 A. The keyhole plasma mode is above 100 A. There is a certain degree of overlap between these current ranges. For example, keyholing can be achieved at 70 A on a 2 mm (0.08 in.) sheet. Equipment is available for welding currents up to 500 A, although a 300 A maximum is typical. Microplasma and medium-current melt-in modes are used for material up to 3 mm (0.12, or

1 , in.) thick, whereas the keyhole plasma mode is used for 8

greater thicknesses and higher travel speeds. In addition to operating in a continuous and steady direct current electrode negative (DCEN) mode, the PAW process can be carried out using DCEN pulsed current, as well as in the variable polarity mode, which uses both direct current electrode positive (DCEP) and electrode negative polarity switching. The pulsed current mode (both DCEN and DCEN/DCEP) is most often used when current levels (typically, above 100 A) are employed for keyhole plasma welding. Pulsing the current widens the tolerance region of acceptance welding parameters, primarily by further stabilizing the formation of the keyhole itself (Fig. 2).

FIG. 2 TOLERANCE TO VARIATION IN WELDING CURRENT AND PLASMA GAS FLOW RATE IN PULSED AND CONTINUOUS CURRENT KEYHOLE WELDING; BOUNDARIES SHOW THE WELDING PARAMETER COMBINATIONS AT WHICH SPECIFIC DEFECTS ARE LIKELY TO OCCUR. WELDING PARAMETERS; NOZZLE BORE, 2.36 MM (0.0929 IN.); ELECTRODE DIAMETER, 4.8 MM (0.19 IN.)

The electrode positive component of the variable polarity plasma arc (VPPA) welding process promotes cathode etching of the tenacious surface oxide film when welding aluminum alloys, allowing good flow characteristics and consistent bead shape. Pulsing times are typically 20 ms for the electrode negative component and 3 ms for the electrode positive polarity. The VPPA welding process is used very effectively in specialized aerospace applications. The PAW process is generally applied when the high penetration of the keyhole welding mode can be exploited to minimize the number of welding passes and, hence, welding time. The time saved can reduce the direct labor element of the welding operation. At the other end of the scale, the microplasma operating mode is used to weld small, thin-section components (as low as 0.025 mm, or 1 mil, thick), where the high arc constriction and low welding current can be beneficial in controlling heat input and distortion. Advantages and Disadvantages. The advantages of the PAW process are primarily intrinsic to the keyhole mode of operation, because greater thicknesses of metal can be penetrated in a single pass, compared with other processes, such as GTAW. This greater amount of penetration allows a reduced amount of joint preparation. In some materials, for example, a square-grooved butt joint preparation can be used for thicknesses up to 12 mm (0.5 in.). The process can produce high weld integrity (similar to GTAW) while minimizing weld passes and, hence, welding times and labor costs. The columnar shape of the arc results in a greater tolerance to variations in torch standoff distance, when compared with the conical arc shape of a GTAW arc. The tungsten electrode used in the PAW process is protected from contamination by the constricting nozzle (Fig. 1). The longer arc length allows better viewing of the weld pool, which is important in manual welding.

Disadvantages include the greater capital equipment cost, when compared with its main rival, the GTAW process. Although high arc constriction achieves higher penetration, it also reduces the tolerance of the process to joint gaps and misalignment, when compared with the broader, conical arc of the GTAW process. The greater complexity of the PAW torch design and the greater number of parts requires more scheduled maintenance. The accurate set-back of the electrode tip, with respect to the nozzle orifice, is required to maintain consistent results. However, this task is facilitated by a general-purpose tool designed for nozzle removal and replacement and for electrode set-back adjustment. Plasma Arc Welding Ian D. Harris, Edison Welding Institute

Equipment A basic PAW system consists of a power source, a plasma control console, a water cooler, a welding torch, and a gas supply system for the plasma and shielding gases (Fig. 3).

FIG. 3 TYPICAL EQUIPMENT FOR PLASMA ARC WELDING

The power source, which supplies the main power for the welding system, is usually supplemented with a sequence controller and control console. The sequence controller sequences the timing of gas flow, arc initiation, main welding current control, and any up-slope and downslope parameters. In its simplest form, the plasma control console controls the gas flow for plasma and shielding gases from separate flow-meters and incorporates the high-frequency pilot arc initiation circuit. The welding torch can be manual or mechanized and is water cooled to avoid torch overheating and to maximize component life. In most PAW installations, plasma and shielding gases are supplied from separate gas cylinders, although bulk gas can readily be used. The gas supply is usually routed through the plasma control console, where the individual flow rates are set by the operator. The power source should be of a constant-current design. Transistorized power sources are most common, although

inverter power supplies are also available. It should have a minimum open-circuit voltage of 80 V to ensure the reliable initiation and transfer of the main arc current. The power source can be adjusted for welding current and it should have the capability to adjust the up slope and down slope of the current. It may be equipped with thumbwheels or potentiometers to select the parameters for pulse current operation, that is, peak and background current levels, as well as peak and background times. Welding Torches. Like those of the other arc welding processes, PAW torches are available in a range of sizes for

different power ratings and in manual and mechanized versions. The design principles are the same in each case. A tungsten alloy electrode is held in a collet within the torch body. To avoid one of the most common defects in plasma torches, it is critical to hold concentricity between the tungsten electrode and the orifice in the design and manufacture of the torch. The electrode assembly is set inside a plenum chamber and the plasma gas is supplied to this chamber. A

threaded copper alloy nozzle forms the front of this chamber and contains the nozzle orifice that is used to constrict the plasma arc. A shielding gas nozzle, usually of an insulating ceramic material, is threaded onto the front end of the torch and surrounds the constricting nozzle, creating an annulus through which the shielding gas is supplied. The torch is connected electrically to the power source and the electrode forms the negative pole of the circuit for dc welding. The gas hoses that supply the plasma and shielding gases and the water hoses that supply and remove water from the torch are all connected to the torch body or handle. These hoses are enclosed in a flexible sheath that extends from the torch to the components of the welding system. Most constricting nozzles have a single orifice in the center. However, multiple-nozzle orifices can be used with higherpower torches to achieve further arc constriction. The most common version of this type of nozzle has a central orifice flanked by a smaller orifice on each side. The common centerline of the three orifices is arranged at 90° to the weld line during operation. Electrodes. The nonconsummable electrode employed is usually a 2% thoriated tungsten electrode, that is, tungsten with

2% thorium oxide. The electrode specification is covered in AWS A5.12-92. The electrode size is selected according to the welding current level that will be used. The electrode is ground with a tapered point on the end, the angle of which depends on the selected welding current level. Electrode sizes and tip angles for a range of welding current levels are shown in Table 1.

TABLE 1 MAXIMUM CURRENT FOR PLASMA WELDING WITH SELECTED ELECTRODE DIAMETER, VERTEX ANGLE, AND NOZZLE BORE DIAMETER MAXIMUM CURRENT, A

ELECTRODE DIAM

mm MICROPLASMA TORCH RATING, 20 A 5 5 1.0 10 20 MEDIUM CURRENT TORCH RATING, 100 A 30 2.4 50 75 100 TORCH RATING, 200 A 50 4.8 100 160 200 TORCH RATING, 400 A 180 3.2 200

VERTEX ANGLE, DEGREES

in.

PLASMA(A) NOZZLE BORE DIAM

SHIELDING(B) SHROUD FLOW RATE DIAM

FLOW RATE

mm

in.

l/min gal/min mm in.

l/min gal/min

0.04

15

0.8 0.8 1.0

0.03 0.03 0.04

0.2 0.3 0.5

0.05 0.08 0.13

8

0.32 4-7

1-1.8

0.10

30

0.79 1.17 1.57 2.06

0.03 0.05 0.06 0.08

0.47 0.71 0.94 1.18

0.12 0.19 0.25 0.31

12

0.48 4-7

1-1.8

0.19

30

1.17 1.57 2.36 3.20

0.05 0.06 0.09 0.13

0.71 0.94 1.42 1.65

0.19 0.25 0.38 0.44

17

0.68 4-12

1-3.2

0.13

60(C)

2.82

0.11

2.4

0.63

18

0.72 2035

5.3-9.2

2.82(C) 0.11(C) 2.5

0.66

HIGH CURRENT TORCH RATING, 400 A 250 4.8

0.19

300 350 (A) (B) (C) (D)

60(C)

3.45(D) 0.14

3.0

0.79

3.45(D) 0.14(D) 3.5 3.96(D) 0.16(D) 4.1

0.92 1.08

...

...

2035

5.3-9.2

ARGON PLASMA GAS. ARGON AND AR-5H2 SHIELDING GAS. ELECTRODE TIP BLUNTED TO 1 MM (0.04 IN.) DIAMETER. MULTIPORT NOZZLE

The new types of tungsten electrodes, which contain oxides of rare-earth elements in place of the thorium oxide, can also be used. These electrodes have been shown to have greater tip life. However, they are more expensive and their usefulness in plasma arc welding may be limited because of the high level of protection provided by the nozzle. In lowcurrent microplasma welding applications, their better emissivity provides easier arc transfer and better overall performance. If a finer wire is required, to fit the joint or to avoid undercut at the weld toes, then wires suitable for gas-shielded arc welding (for example, those used for GTAW) should be employed. The appropriate AWS specification series is AWS A5.XX. Plasma (Orifice) and Shielding Gases. The plasma gas is used to generate the arc, whereas the shielding gas is used to provide the weld pool with supplementary shielding from atmospheric contamination while it solidifies and cools. The plasma gas is almost always argon. Gas properties affect both weld shape and quality. Flow rates, particularly of the plasma gas, are also important, because they control the extent of plasma constriction. The flow rate can vary from 0.1 L/min (0.026 gal/min) for microplasma welding up to 10 L/min (2.6 gal/min) for keyhole plasma welding. Considerable care is needed to regulate the gas flow rate if keyhole closure is required, because the flow rate must be sloped out to 1 to 2 L/min (0.26 to 0.52 gal/min) within about 1 s. Gas flow control is best achieved by electronic means.

The design and current rating of welding torches are based on argon plasma gas. Argon provides effective shielding, being heavier than air, and is cheaper than helium. Shielding gas selection is based on the type of base metal (Table 2).

TABLE 2 PLASMA AND SHIELDING GAS COMPOSITIONS

MATERIAL MILD STEEL

PLASMA GAS SHIELDING GAS ARGON ARGON ARGON-2-5% H2(A) LOW-ALLOY STEELS ARGON ARGON AUSTENITIC STAINLESS STEEL ARGON ARGON-2-5% H2 HELIUM(A) NICKEL AND NICKEL ALLOY ARGON ARGON ARGON-2-5% H2(A) TITANIUM ARGON ARGON 75HE-25AR(A) ALUMINUM AND ALUMINUM ALLOYS ARGON ARGON HELIUM(A) COPPER AND COPPER ALLOYS ARGON ARGON 75HE-25AR(A) (A)

ALSO USED

Helium and argon-helium mixtures can be used as shielding gases to increase the thermal conductivity of the gas and, hence, the heating effect on the weld pool. Helium results in wider weld pools than argon, because it produces a higher arc voltage. Hydrogen additions to argon shielding gas tends to promote slightly narrower weld pools through arc constriction and achieves a very clean weld pool appearance, because it is a reducing gas. Although helium and hydrogen can be added to argon in the shielding gas to give higher heat input, the use of gases with higher heat contents in the plasma gas can result in torch overheating and potential damage. Plasma Arc Welding Ian D. Harris, Edison Welding Institute

Applications Material Types. The PAW process is commonly used to weld stainless steels in a wide range of thicknesses. The process

can also be used with carbon and alloy steels, aluminum alloys, titanium alloys, copper and nickel alloys, and more specialized materials, such as zirconium and tantalum. The thicknesses that can be welded in a single pass range from 0.025 mm (1 mil) for microplasma applications to 12.5 mm (0.5 in.) for the VPPA welding of aluminum. Direct-current pulsing can be used on most materials. The PAW process is often carried out in an autogenous mode, that is, without filler wire. When edge beveling is used, a filler wire is required to complete the joint. A filler wire can also be used with the keyhole mode of operation to avoid undercut at high welding speeds. Wire composition depends on that of the parent materials in the joint. The same continuous-wound wire that is used in GTAW operations is suitable. The industries that use the PAW process can be categorized as those that weld thin-section sheet using microplasma or medium-current plasma welding and those that weld plate using keyhole plasma welding.

A wide range of small devices and assemblies made from thin stainless steel sheet, including bellows assemblies and associated fittings, are welded using the microplasma operating mode. The narrow weld bead that can be produced provides sheet-metal fabrications with a good cosmetic appearance. Furthermore, the high welding speed that can be achieved, coupled with the good tolerance to stand-off variations resulting from the columnar nature of the arc, makes the process attractive for high-volume production work. Microplasma, as well as medium-current plasma modes, can be used to spot weld guide wires and lamp filaments, as well as in other applications that require highly repetitive autogenous welds. This type of application allows a user to limit the number of high-frequency arc starts that would be required with tungsten-inert gas welding. Keyhole plasma welding is extensively used to weld stainless steel pipe and tankage. The process is applied to individual strakes from plate to make stainless steel vessels in the food and chemical processing industries. Circumferential welding of strakes also can be used to create these products. Longitudinal seam welds in stainless steel pipe with wall thickness of [ges]3 mm ([ges]0.12 in.) are ideally suited to keyhole plasma welding, because joint preparation is minimized and single-pass welding can be consistently achieved without the use of weld-backing devices. Pipes with wall thickness above 5 to 6 mm (0.20 to 0.24 in.) employ the keyhole mode of operation for the root pass. Depending on the material type and wall thickness, melt-mode PAW, GTAW, gas-metal arc welding (GMAW), or submerged arc welding (SAW) are used to complete the joint. High-alloy composition piping is similarly manufactured. The manufacturing of stainless steel tube from strip was one of the first applications of the PAW process. Because the process can reliably produce full-penetration welds without the use of backing, it is extensively used on tube mills, because a lack of access precludes welding from the inside. The VPPA welding process has been very effective for welding large space shuttle orbiter sections of aluminum alloys, particularly the external fuel tank. The cathodic cleaning action of the electrode positive portion of the current cycle breaks up tenacious surface oxide film on the aluminum. The process can be used on thicknesses up to 12.5 mm (0.5 in.) with square-grooved butt joints, and it can be used for the root pass on thicker sections with edge beveling. It is anticipated that this technique will be extensively employed in building aluminum alloys structures in a modular space station. Plasma Arc Welding

Ian D. Harris, Edison Welding Institute

Typical Components and Joints The most common joint configuration used with the PAW process is a butt joint. The microplasma mode is used with overlapped butt (micro-lap) joints and with joints that have integral weld metal as a result of flanged, butted edges on very thin metals (Table 3). Corner joints with edge welds are also commonly welded using the microplasma and mediumcurrent modes.

TABLE 3 JOINT CONFIGURATIONS FOR PLASMA WELDING SHEET AND TUBULAR COMPONENTS

Because the keyhole operating mode fully penetrates the workpiece, it is used exclusively on square-grooved butt joints and single-V joints with a root face (Table 3). For square-grooved butt joint preparation, the thickness that can be welded in a single pass depends on the fluid flow characteristics of the workpiece material for a given heat input from the plasma torch. Thus, alloys of titanium and zirconium can be butt welded with square-grooved preparation at greater thicknesses than steels and stainless steels. Generally, it is an industry-accepted practice to use square-grooved preparation without edge beveling for stainless steels up to 6 mm (0.24 in.) in thickness. A 4 or 5 mm (0.16 or 0.20 in.) root face and a single-V preparation, usually with a 60° included angle, are used on steels thicker than 6 mm (0.24 in.) (Table 4). The root is welded in the keyhole mode and the rest of the joint is completed with a filler wire addition by melt-mode PAW, GTAW, GMAW, or SAW, depending on the material type, joint volume to be filled, and the mechanical property requirements of the joint.

TABLE 4 KEYHOLE PLASMA WELDING CONDITIONS FOR PIPE PREFABRICATION IN TYPE 304L/316L STAINLESS STEEL

(a) Welded thickness, 5 mm (0.20 in.). (b) Welded thickness, 6 mm (0.24 in.)

In the PAW process, fixturing for part fit-up is more critical than it is for the GTAW process, primarily because of the narrower, more-constricted arc in PAW. The backing bar design and shielding gas techniques employed for GTAW are appropriate for microplasma and medium-current modes. During keyhole welding, however, the arc passes right through the workpiece, forming an efflux plasma (Fig. 1) that normally extends 10 mm (0.4 in.) below the back face of the joint. This characteristic must be accommodated when designing backing systems. Failure to allow sufficient gap can result in turbulence in the efflux plasma, causing weld pool disturbance and porosity. When welding longitudinal seams in tube or pipe, the weld can be started and stopped on run-on and run-off tabs, precluding the need for keyhole closure. However, when using the keyhole mode on circumferential seams of vessels or on butt joints in tubes or pipes, the keyhole will overlap the start of the weld to produce a complete joint. Keyhole closure without porosity has been an area of difficulty, particularly when butt welding tubes, and it represented an important concern for more-widespread use of the process. Modern electronic controls for simultaneous reduction in gas flow and current slope-out provide a more reliable method for keyhole closure. Plasma Arc Welding Ian D. Harris, Edison Welding Institute

Procedures Process Operating Procedure. Welding parameters, such as welding current, arc voltage, travel speed (for mechanized/automatic operation), and plasma and shielding gas flow rates, are set by the procedure and implemented by the welder. Torch parameters include the correct electrode vertex angle and set-back distance, as well as the correct orifice diameter for the welding current level.

The operating sequence for the PAW process was described in the section "Principles of Operation" at the beginning of this article. After the weld pool or keyhole is formed, the torch is traversed across the workpiece at the preset welding speed. Welding is terminated by the down slope of the welding current, with simultaneous sloping of the plasma gas flow rate if keyhole welding and keyhole closure are desired. Slope control for the keyhole mode is shown in Fig. 4.

FIG. 4 TYPICAL SLOPE CONTROL PATTERN FOR WELDING CURRENT AND PLASMA GAS FLOW WHEN STARTING AND CLOSING A KEYHOLE; EXAMPLE IS FOR 9.5 MM (

3 IN.) THICK STEEL 8

Most plasma arc welding is done in a mechanized or automated operation that does not require much operator intervention, except initially to set the parameters and to position the workpiece and torch. When current pulsing and keyhole closure operations are involved, numerous parameters must be set, which requires strict attention to detail, accurate part fit-up, and careful alignment of the torch relative to the joint. Inspection and Weld Quality Control. All of the common nondestructive evaluation (NDE) techniques are applicable

to plasma arc welds. Radiography and ultrasonic inspection are the most common techniques used. The fact that most plasma arc welding is carried out in the keyhole mode means that there is a penetration bead on the root side of the joint. Visual inspection for full penetration, as well as for correct width and profile of both the penetration bead and the weld surface profile, can readily be accomplished. Smooth, even underbeads can be achieved through good procedure development with tolerances improved through current pulsing. Troubleshooting. When troubleshooting a PAW operation, all of the previously defined basic parameters of the welding

schedule should be checked. Welding discontinuities can result from incorrect electrode set-back or from a worn or damaged nozzle orifice. The concentricity of the diameter of the nozzle orifice and the alignment of the electrode and nozzle orifice are very important. Worn nozzles should be replaced. The condition of the tungsten electrode tip should also be checked. An undercut is one of the most common defects in the PAW process. Depending on material type, thickness, and preset welding parameters, a two-sided undercut can occur above certain threshold welding speeds. This can only be eliminated by increasing welding current. A one-sided undercut can result from misalignment of the torch, the electrode/orifice and multiport nozzles, or from the mismatched fit-up of the workpieces. Correct alignment is essential. Plasma Arc Welding Ian D. Harris, Edison Welding Institute

Personnel Requirements The skill level and training requirements for the PAW process are generally similar to those for the GTAW process. When manual welding, the columnar nature of the plasma arc allows greater variation of the torch-to-workpiece distance without altering the size and shape of the weld pool. The fact that the electrode is protected from contamination by the nozzle means that tungsten contamination and tungsten inclusions in the weld metal are more readily avoided than they are in the GTAW process. From the viewpoint of an operator, disadvantages include the greater complexity of the torch,

in terms of nozzle orifice and electrode set-back considerations, and the correct setting of two gases (plasma and shielding gas). When mechanized welding is carried out, which is usually the case for the keyhole mode, the operator is required to set a greater number of parameters (compared to GTAW), particularly if current pulsing and keyhole closure operations are required. Specific operator training is required for pulsed keyhole operation, and particularly for VPPA operations, which are usually carried out in the vertical position, welding upward, and thus require a high level of operator competence. Computerization of welding parameter selection is the trend for VPPA welding and other mechanized or automated PAW operations. The health and safety issues related to the PAW process are very similar to those of other gas-shielded arc welding processes, especially GTAW. These include electrical shock; electromagnetic radiation hazards, particularly ultraviolet radiation; burns from hot metal parts; and welding fumes and gases, including ozone. The volume of welding fumes produced is low and is similar to that produced by the GTAW process. Hexavalent chromium and ozone are concerns when welding stainless steels and aluminum alloys, respectively. Like the GTAW process, the low level of general welding fumes associated with PAW results in a comparatively higher level of ozone. However, the rapid decay of ozone within a short distance of the arc, coupled with the high degree of mechanization and automation typical of the keyhole plasma welding operation, means that operator exposure is generally very low. The microplasma mode, which is more commonly used manually, employs such low current levels that fume and ozone levels are very low. Carbon Arc Welding Lance Soisson, Welding Consultants, Inc.

Introduction CARBON ARC WELDING (CAW) utilizes what is considered to be a nonconsumable electrode, made of carbon or graphite, to establish an arc between itself and either the workpiece or another carbon electrode. However, this electrode erodes fairly quickly and generates carbon monoxide (CO) gas that partially replaces the air around the arc, thereby providing the molten weld with some protection. The CAW process, which uses either single or twin electrodes, most closely resembles gas-tungsten arc welding (GTAW), where the arc is used only as a source of heat. The single-electrode arrangement usually operates with direct current (dc), electrode negative (straight polarity), using most dc power supplies. The twin-electrode arrangement usually operates with alternating current (ac), generally with small ac power supplies. Figure 1 shows typical configurations for both procedures.

FIG. 1 TYPICAL ARRANGEMENTS FOR SINGLE-ELECTRODE AND TWIN-ELECTRODE CARBON ARC WELDING

Although the CAW process has been almost completely replaced by the newer processes used in the welding industry, it is still used in certain applications. The process does produce adequate welds in thin-sheet steel, but should be used with caution in any critical application because it provides only limited shielding from the atmosphere. Carbon Arc Welding Lance Soisson, Welding Consultants, Inc.

Operation CAW electrodes can be either carbon or graphite, although baked carbon electrodes are most commonly used because they last longer and are more readily available. Electrode sizes range from 3.2 to 22 mm (

1 7 to in.). The typical torch is 8 8

air cooled for the smaller electrodes, but water cooled for the larger ones. The recommended current ranges for both electrode types are given in Table 1.

TABLE 1 RECOMMENDED CURRENT RANGES FOR CARBON AND GRAPHITE ELECTRODES

ELECTRODE DIAMETER CURRENT, A mm in. CARBON GRAPHITE SINGLE ELECTRODE 1 3.2 15-30 15-35 4.8 6.4 7.9 9.5 11.1 12.7 15.9 19 22.2

8 3 16 1 4 5 16 3 8 7 16 1 2 5 8 3 4 7 8

TWIN ELECTRODES 1 6.4 7.9 9.5

4 5 16 3 8

25-55

25-60

50-85

50-90

75-115

80-125

100-150

110-165

125-185

140-210

150-225

170-260

200-310

230-370

250-400

290-490

300-500

400-750

55

...

75

...

95-120

...

In the single-electrode operation, the carbon electrode is prepared to a long tapered point that is one half the

electrode diameter (see Fig. 1). The arc is established by "scratch starting," that is, by bringing the electrode into contact with the workpiece and immediately withdrawing it to the correct length for welding. The generally accepted arc length is maintained between 6.4 and 9.5 mm (

1 3 and in.). Holding the length too short can result in carburization of the base 4 8

metal, creating potentially brittle welds. This is particularly true when welding without filler metal. When the joint requires filler metal, the same technique that is used in GTAW or oxyfuel gas welding (OFW) should be employed. The welding rod should be fed into the weld pool with one hand while manipulating the arc with the other. If the arc is broken for any reason, then restart should not occur directly at the hot metal where welding stopped, because a hard spot could result. Rather, a restart should occur in front or to one side of the weld, followed by a quick return to the point where welding stopped. Care should be taken to either incorporate the restart area into the weld or to grind the arc strike after welding is completed. The arc strike areas can be very hard and can cause cracking if they are not fused into the weld or ground out. In the twin-electrode operation, the workpiece is not part of the electrical circuit. The heat for welding is produced

by creating an arc between the two electrodes. The electrode holder has two adjustable arms that grip the carbon electrodes at the proper length to maintain the arc. The operator maintains a constant distance between the carbon electrodes by adjusting them as they are consumed. An experienced welder can make these adjustments while still welding or brazing. Small ac machines are normally used in the twin-electrode process because the electrodes will wear at an equal rate. Direct current can be used as long as the positive-side electrode is one size larger than the negative electrode. The larger electrode compensates for the faster consumption rate on the positive side of the direct current. 1 3 to in.) and are copper coated for better current flow. The heat must be 4 8 1 maintained so that the copper does not melt more than 13 mm ( in.) back from the electrode tip. Only enough heat to 2

Electrodes range in size from 6.4 to 9.5 mm (

cause the filler metal to flow easily should be used (Table 1). High heat causes excessive electrode consumption and could cause carburization of the weld metal. Carbon Arc Welding Lance Soisson, Welding Consultants, Inc.

Applications The single-electrode carbon arc process is primarily used on steel, cast iron, and copper, although it can be used on most ferrous and nonferrous materials. With steels, it is principally applied to make outside corner welds on thinner-gage materials, where no filler metal is used. A good fit-up is required, and a fluxing agent is often used to promote better welds. The resulting welds arc smoother and, frequently, more economical than similar welds created by the shielded metal arc welding (SMAW) process. Another application is the welding of galvanized steel using a silicon bronze filler metal. In this instance, it is important to maintain a short arc and to direct the heat to the filler metal to minimize damage to the galvanic coating. Cast iron can be welded using the carbon arc process, but it is more often used in brazing. The casting must be preheated to approximately 650 °C (1200 °F) and then slowly cooled to reduce the possibility of cracking and to produce machinable welds. The process is also used on copper, because of the high heat capabilities of the larger electrodes. Direct current, electrode negative, should always be used when welding or brazing copper. The workpiece should be preheated to a temperature between 150 and 650 °C (300 and 1200 °F), or at least locally preheated in the weld area. Holding a long arc will allow the carbon to combine with oxygen to form the CO that somewhat shields the weld.

Twin carbon arc welding is used primarily as a maintenance tool in small shops, farms, or homes. It is most commonly applied in brazing or soldering operations, although it can be used to weld, preheat, or postheat smaller parts. Carbon Arc Welding Lance Soisson, Welding Consultants, Inc.

Selected References

• H.B. CARY, MODERN WELDING TECHNOLOGY, 2ND ED., PRENTICE-HALL, 1989 • WELDING HANDBOOK, 8TH ED., VOL 2, WELDING PROCESSES, AWS, 1991 Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

Introduction SUBMERGED ARC WELDING (SAW) is an arc welding process in which the arc is concealed by a blanket of granular and fusible flux (Ref 1, 2, 3). Heat for SAW is generated by an arc between a bare, solid-metal (or cored) consumablewire or strip electrode and the workpiece. The arc is maintained in a cavity of molten flux or slag, which refines the weld metal and protects it from atmospheric contamination. Alloy ingredients in the flux may be present to enhance the mechanical properties and crack resistance of the weld deposit.

References

1. R.L. O'BRIEN, WELDING HANDBOOK, VOL 11, AWS, 1991, P 191-232 2. D.L. OLSON ET AL., SUBMERGED ARC WELDING, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 114-152 3. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 6.3-1 TO 6.3-24 Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

Principles of Operation Figure 1 shows a typical setup for automatic SAW. A continuous electrode is being fed into the joint by mechanically powered drive rolls. A layer of granular flux, just deep enough to prevent flash through, is being deposited in front of the arc. Electrical current, which produces the arc, is supplied to the electrode through the contact tube. The current can be direct current (dc) with electrode positive (reverse polarity), with electrode negative (straight polarity), or alternating current (ac). Figure 2 shows the melting and solidification sequence of SAW. After welding is completed and the weld metal has solidified, the unfused flux and slag are removed. The unfused flux may be screened and reused. The solidified slag may be collected, crushed, resized, and blended back into new flux. Recrushed slag and blends of recrushed slag with unused (virgin) flux are chemically different from new flux. Blends of recrushed slag may be classified as a welding flux, but cannot be considered the same as the original virgin flux (Ref 4).

FIG. 1 SCHEMATIC SHOWING KEY COMPONENTS OF AUTOMATIC SUBMERGED ARC WELDING SETUP

FIG. 2 SCHEMATIC SHOWING TYPICAL WELD POOL DYNAMICS OF A SUBMERGED ARC WELD

Submerged arc welding is adaptable to both semiautomatic and fully automatic operation, although the latter, because of its inherent advantages, is more popular (see the section "Advantages of Submerged Arc Welding" in this article). In semiautomatic welding, the welder controls the travel speed, direction, and placement of the weld. A semiautomatic welding gun is designed to transport the flux and wire to the operator, who welds by dragging the gun along the weld joint. Semiautomatic electrode diameters are usually less than 2.4 mm (

3 in.) to provide sufficient flexibility and 32

feedability in the gun assembly. Manually guiding the gun over the joint requires skill because the joint is obscured from view by the flux layer. In automatic SAW, travel speed and direction are controlled mechanically. Flux may be

automatically deposited in front of the arc, while the unfused flux may be picked by a vacuum recovery system behind the arc. To increase deposition rate or welding speed, more than one wire can be fed simultaneously into the same weld pool. For example, in the twin arc process, two electrodes are fed into the same weld pool while sharing a common power source and contact tip. In tandem arc SAW, multiple electrodes are arranged with one in front of the other. Each electrode has an independent power supply and contact tip. The spacing, configuration, and electrical nature of the electrodes may be arranged to optimize welding speed and bead shape (Ref 5). Advantages of submerged arc welding include the following: •

• • • •

• • • •

THE ARC IS UNDER A BLANKET OF FLUX, WHICH VIRTUALLY ELIMINATES ARC FLASH, SPATTER, AND FUME (THUS MAKING THE PROCESS ATTRACTIVE FROM AN ENVIRONMENTAL STANDPOINT). HIGH CURRENT DENSITIES INCREASE PENETRATION AND DECREASE THE NEED FOR EDGE PREPARATION. HIGH DEPOSITION RATES AND WELDING SPEEDS ARE POSSIBLE. COST PER UNIT LENGTH OF JOINT IS RELATIVELY LOW. THE FLUX ACTS AS A SCAVENGER AND DEOXIDIZER TO REMOVE CONTAMINANTS SUCH AS OXYGEN, NITROGEN, AND SULFUR FROM THE MOLTEN WELD POOL. THIS HELPS TO PRODUCE SOUND WELDS WITH EXCELLENT MECHANICAL PROPERTIES. LOW-HYDROGEN WELD DEPOSITS CAN BE PRODUCED. THE SHIELDING PROVIDED BY THE FLUX IS SUBSTANTIAL AND IS NOT SENSITIVE TO WIND AS IN SHIELDED METAL ARC WELDING AND GAS METAL ARC WELDING. MINIMAL WELDER TRAINING IS REQUIRED (THUS, RELATIVELY UNSKILLED WELDERS CAN BE EMPLOYED). THE SLAG CAN BE COLLECTED, REGROUND, AND SIZED FOR MIXING BACK INTO NEW FLUX AS PRESCRIBED BY MANUFACTURERS AND QUALIFIED PROCEDURES.

Limitations of submerged arc welding include the following: • • • •

THE INITIAL COST OF WIRE FEEDER, POWER SUPPLY, CONTROLS, AND FLUXHANDLING EQUIPMENT IS HIGH. THE WELD JOINT NEEDS TO BE PLACED IN THE FLAT OR HORIZONTAL POSITION TO KEEP THE FLUX POSITIONED IN THE JOINT. THE SLAG MUST BE REMOVED BEFORE SUBSEQUENT PASSES CAN BE DEPOSITED. BECAUSE OF THE HIGH HEAT INPUT, SAW IS MOST COMMONLY USED TO JOIN STEELS MORE THAN 6.4 MM (

1 4

IN.) THICK.

References cited in this section

4. "SPECIFICATION FOR CARBON STEEL ELECTRODES AND FLUXES FOR SUBMERGED ARC WELDING," A5.17-89, AWS, 1989, P 18 5. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 3.2-3, 6.3-21 TO 6.3-22, 6.3-49 TO 6.3-58 Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

Process Applications

If a steel is suitable for welding with gas-metal arc welding, flux-cored arc welding, shielded metal arc welding, or gastungsten arc welding, procedures can be developed to weld the steel with SAW (Ref 6). The main limitations of SAW are plate thickness and position. Because SAW is a high heat input and high deposition rate process, it is generally used to weld thicker-section steels. Although the welding of 1.6 mm ( over 6.4 mm (

1 in.) thick steel is possible, most SAW is done on plate 16

1 in.) in thickness. 4

Types of Metals. Submerged arc welding is most commonly used to join plain carbon steels. Alloy steels can be readily

welded with SAW if care is taken to limit the heat input as required to prevent damage to the heat-affected zone (HAZ). Low-heat-input procedures are available for welding alloy steels and heat-treated steels to prevent grain coarsening and cracking in the HAZ. Maintaining proper preheat and interpass temperature is also important when welding alloy steels to prevent weld metal and HAZ cracking and to develop the required mechanical properties in the weld deposit. Submerged arc welding can be used to join stainless steels and nonferrous alloys. It is also commonly used to produce a stainless or nonferrous overlay on top of a base metal. Joint Configurations. The most common weld deposits made with SAW are groove, fillet, and plug welds, and

surfacing deposits. For groove welds, the characteristic deep penetration of SAW plays a role in joint selection. Plate up 5 1 in.) thick can be completely welded from one side using a square butt joint with a 0.8 mm ( in.) root 8 32 5 opening. Beveled joints should be used for multiple-pass welds on plate thicknesses greater than 15.9 mm ( in.). 8

to 15.9 mm (

Backing bars should be used to prevent loss of flux, to prevent melt through, and to ensure full-penetration weldments on one-sided joints. Although groove welds are generally made in the flat position, it is possible to develop special procedures to weld horizontally. Single-pass fillet welds with throat sizes up to 7.9 mm (

5 in.) thick and multiple-pass fillet welds are usually made in 16

the horizontal position. Single-pass fillet welds over 7.9 mm (5/16 in.) thick are usually made in the flat position. Submerged arc plug welds are made in the flat position by puddling the electrode into the center of the hole until the weld is complete. Surfacing welds are made with both wire and strip electrode. Surfacing is done to rebuild worn parts with a wear-resistant material or to overlay a plain steel part with stainless steel or other alloys (Ref 7, 8, 9). To prevent porosity, the surface to be welded should be clean and free of grease, oil, paints, moisture, and oxides. All slag from tack welds or previous layers should be removed. Tack welds should be positioned so that the submerged arc weld completely melts out the tack. The workpiece should be firmly clamped to minimize distortion and the need for tack welds. Industrial Uses for SAW. Because SAW is used to join thick steel sections, it is primarily used for shipbuilding, pipe

fabrication, pressure vessels, and structural components for bridges and buildings. Other than joining, SAW is used to build up parts and overlay with stainless or wear-resistant steel (for example, rolls for continuous casting steel, pressure vessels, rail car wheels, and equipment for mining, mineral processing, construction, and agriculture).

References cited in this section

6. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 6.3-26 TO 6.3-73 7. R.S. BROWN ET AL., ARC WELDING OF STAINLESS STEELS, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 327-329 8. K.C. ANTONY ET AL., HARDFACING, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 784 9. D.R. THOMAS ET AL., WELD OVERLAYS, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 804-819

Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

Equipment and Supplies The power supplied to the contact tip may be direct current or alternating current. In the United States, dc power is most common for welding applications that require less than 1000 A. In Europe and the Far East, however, ac is used extensively, even in applications requiring less than 1000 A. Currents over 1000 A dc tend to create arc blow problems. Alternating current is most commonly used for high-current applications, for applications where arc blow may be a problem, and in multiwire applications. Direct-current power supplies may be either transformer-rectifiers or motor/engine generators that produce either constant voltage or constant-current power. Alternating-current power sources are most commonly transformers. To produce a submerged arc weld, both flux and electrode are consumed. Each flux and electrode combination, along with the variation of base material and process parameters, will produce a unique weld deposit. Because the integrity of the weld deposit depends on these parameters, specific fluxes and electrodes must be used in combination to optimize the weld metal properties. Power Sources A constant-voltage power supply is self-regulating, so it can be used with a constant-speed wire feeder. No voltage or current sensing is required. The current is controlled by the wire diameter, the electrical stickout, and the wire-feed speed, while the voltage is controlled by the power supply. Constant-voltage dc power is the best choice for the high-speed welding of thin steel. Unlike constant voltage, constant-current power supplies are not self-regulating, so they must be used with voltagesensing variable-wire-feed speed controls. A constant-current wire feeder monitors arc voltage and adjusts the wire-feed speed in response to changes in the arc voltage. The wire-feed speed control attempts to maintain a constant arc length, while the power supply controls the arc current. The constant-current output of a conventional ac machine varies with time like a sine wave dropping through zero with each polarity reversal. The voltage associated with the current is approximately a square wave. Because the power output is zero with each polarity change, an open-circuit voltage greater than 80 V may be required to ensure arc initiation. The constant-current ac machine requires voltage-sensing variable-wire-feed speed controls. On newer, solid-state power supplies, the current and voltage output both approximate square waves, with the instantaneous polarity reversal reducing arc initiation problems. The solid-state power supplies have constant-voltage characteristics that may be used with constant-speed wire-feed controls. Flux Classification Fluxes can be categorized depending on the method of manufacture, the extent to which they can affect the alloy content of the weld deposit, and the effect on weld deposit properties. Classification Relative to Production Method. Based on the manufacturing process, there are two different types of

fluxes: fused and bonded. Fused Fluxes. The raw materials for a fused flux are dry mixed and melted in a furnace. The molten mixture is then

rapidly solidified, crushed, screened, and packaged. Because of their method of manufacture, fused fluxes typically do not contain ferroalloys and deoxidizers. Bonded Fluxes. The powdered ingredients of a bonded flux are dry blended, and then mixed with a binder, usually

potassium or sodium silicate. After pelletizing, the wet flux is dried in an oven or kiln, sized appropriately, and then packaged. The relatively low baking temperatures allow bonded fluxes to contain deoxidizers and ferroalloys. Classification Relative to Effect on Alloy Content of Weld Deposit. Independent of manufacturing method, a given

flux may be described as an active, neutral, or alloy flux, depending on its ability to change the alloy content of the weld deposit. With all submerged arc fluxes, variations in arc voltage and other welding variables will change the ratio of flux

consumed to electrode or weld metal deposited. This ratio is often referred to as the flux-to-wire ratio. Normal flux-towire ratios for SAW are 0.7 to 0.9. An increase in the flux-to-wire ratio may be caused by either an increase in arc voltage or a decrease in the welding current. Likewise, a decrease in the flux-to-wire ratio may be caused by a decrease in arc voltage or an increase in the welding current. How the weld deposit composition changes with voltage (flux-to-wire ratio) provides an additional means of describing a flux. Active fluxes contain controlled amounts of manganese and/or silicon. These alloys are added as ingredients in the flux to provide improved resistance to porosity and weld cracking caused by contaminants such as oxygen, nitrogen, and sulfur on the plate or in the plate composition itself. Active fluxes are primarily used to make single-pass welds. Because active fluxes contain deoxidizers such as manganese and silicon, the alloy in the weld metal will change with the flux-towire ratio. Changes in manganese and silicon content in the weld deposit will affect the strength and impact properties of the weld metal; therefore, the arc voltage must be more tightly controlled when welding with active fluxes than with neutral fluxes. Neutral fluxes contain little or no deoxidizers, and by definition, will not produce any significant change in the weld

metal composition as a result of a large change in arc voltage. Because neutral fluxes contain little or no alloy, they must rely on alloy in the electrode to provide deoxidation. Single-pass welds on oxidized plate may be prone to cracking or porosity due to insufficient alloy content and are usually best welded with active fluxes. Neutral fluxes are primarily used for welding multiple passes on plate exceeding 25 mm (1 in.) in thickness. To quantify the active/neutral behavior of a flux, the Wall neutrality number (N) was developed (Ref 10) and is defined as:

N = 100 [|∆%SI| + |∆%MN|]

(EQ 1)

where ∆%Si is the change in the weight percent of silicon, and ∆%Mn is the change in the weight percent of manganese. To determine the Wall neutrality number, two different weld deposit pads must be made. One weld pad is welded with the same parameters as those specified for the weld test plate for the flux and electrode being used. The other weld pad is welded with the same parameters except that the voltage is increased by 8 V. The ∆%Si and ∆%Mn quantities in Eq 1 are then determined by the difference in the amount of silicon and manganese in the two deposits. A flux electrode combination that produces a Wall neutrality number of 40 or less is considered neutral, while a number of 40 indicates an active flux. Alloy fluxes contain enough alloy as a flux ingredient to produce an alloy weld metal with a carbon steel electrode. They

are also used with alloy and stainless steel wire and strip electrodes. Alloy fluxes find application primarily in welding alloy steels and hardfacing. Because the alloy in the weld deposit is a function of arc voltage (flux-to-wire ratio), it is important to follow the manufacturer's recommended procedure for obtaining proper weld deposit composition. Classification Relative to Basicity Index. Aside from the manufacturing method and the active, neutral, or alloy

behavior of flux types, another common method used to describe submerged arc fluxes is the basicity index, BI. The basicity index is the ratio of strongly bound metallic oxides to weakly bound metallic oxides (Ref 11). The Boniszewski basicity index is defined by:

(EQ 2)

The basicity index is an estimate of the oxygen content in the weld metal and is therefore used to predict weld metal properties. Basic fluxes tend to have lower weld metal oxygen content with good weld metal toughness, while acidic fluxes tend to produce higher weld metal oxygen content and coarser microstructure with a lower resistance to cleavage. Fluxes with a basicity index greater than 1.5 are considered basic; fluxes below 1.0 are considered acidic. Acid fluxes are typically preferred for single-pass welding because of their superior operating and bead-wetting

characteristics. In addition, acid fluxes usually have more resistance than basic fluxes to porosity caused by plate contamination by oil, rust, and mill scale.

Basic fluxes tend to have better impact properties than acid fluxes. This advantage is particularly evident on large

multiple-pass welds. Highly basic fluxes generally produce weld metal with very good impact properties in large multiple-pass weldments. Basic fluxes tend to exhibit poorer welding characteristics than acid fluxes on single-pass welds. Their use should be limited to large multipass weldments where good weld metal notch toughness is required. Selection of Electrodes Relative to Steel and Alloy Type There are three different types of consumable electrodes: solid wire, cored wire, and strip. Electrodes for SAW are available for welding carbon steel, low-alloy steel, stainless steel, and nickel-base alloys. Wire diameters vary from 1.6 to 6.4 mm (

1 1 to in.). Carbon steel wire usually has a light copper coating, which protects the wire from corrosion and 16 4

provides good electrical contact in the welding tip. Cored electrodes were developed to provide a low-cost method of increasing the number of alloys that can be welded with SAW. Cored electrodes typically consist of a mild steel tube with alloying elements in the center. Compositions of cored electrodes vary from mild and low-alloy steels through tool steels to stainless steels and high-alloy austenitic steels. Strip electrodes are generally used for welding overlays and for hardfacing applications. The advantages of strip are wide weld beads, low penetration, low dilution, and high deposition rates. The equipment is very similar to SAW with wire except that special drive rolls are needed to feed the strip, and sometimes auxiliary magnetic fields are set up to improve bead shape and tie-in. Strip thicknesses vary from 0.5 to 1.0 mm (0.02 to 0.04 in.), and widths vary from about 25 to 100 mm (1 to 4 in.). Plain Carbon Steels. The requirements for the classification of electrodes and fluxes for welding carbon steels are

detailed in Ref 12. Compositional requirements for solid electrodes are given in Table 1. Requirements for composite electrodes are defined in Table 2. Classification of solid electrodes is based on the chemical composition of the electrode. Classification of composite electrodes is based on weld deposit composition. Flux classification is based on the soundness and mechanical properties of the weld metal produced with a particular flux and electrode combination. The classification system used to describe flux-electrode combinations for mild steel welding is defined in Fig. 3. Interpretation of the classification system code is illustrated in the flux-electrode combination examples that follow.

TABLE 1 COMPOSITION OF SOLID-CARBON STEEL ELECTRODES ELECTRODE AWS CLASSIFICATION

UNS NO.

COMPOSITION, WT%(A)(B) C Mn Si

LOW-MANGANESE ELECTRODES EL8 K01008 0.10 0.25/0.60 EL8K K01009 0.10 0.25/0.60 EL12 K01012 0.04/0.14 0.25/0.60 MEDIUM-MANGANESE ELECTRODES EM12 K01112 0.06/0.15 0.80/1.25 EM12K K11113 0.05/0.15 0.80/1.25 EM13K K01313 0.06/0.16 0.90/1.40 EM14K K01314 0.06/0.19 0.90/1.40 EM15K K01515 0.10/0.20 0.80/1.25 HIGH-MANGANESE ELECTRODES EH11K K11140 0.07/0.15 1.40/1.85 EH12K K01213 0.06/0.15 1.50/2.00 EH14 K11585 0.10/0.20 1.70/2.20

S

P

Cu(c)

Ti

0.07 0.030 0.030 0.35 . . . 0.10/0.25 0.030 0.030 0.35 . . . 0.10 0.030 0.030 0.35 . . . 0.10 0.10/0.35 0.35/0.75 0.35/0.75 0.10/0.35

0.030 0.030 0.030 0.025 0.030

0.030 0.030 0.030 0.025 0.030

0.35 0.35 0.35 0.35 0.35

... ... ... 0.03-0.17 ...

0.80/1.15 0.030 0.030 0.35 . . . 0.25/0.65 0.025 0.025 0.35 . . . 0.10 0.030 0.030 0.35 . . .

Source: Ref 12

(A) THE FILLER METAL SHALL BE ANALYZED FOR THE SPECIFIC ELEMENTS FOR WHICH VALUES ARE SHOWN IN THIS TABLE. IF THE PRESENCE OF OTHER ELEMENTS IS INDICATED, IN THE COURSE OF THIS WORK, THE AMOUNT OF THOSE ELEMENTS SHALL

BE DETERMINED TO ENSURE THAT THEIR TOTAL (EXCLUDING IRON) DOES NOT EXCEED 0.50%. (B) SINGLE VALUES ARE MAXIMUM. (C) COPPER LIMIT INCLUDES ANY COPPER COATING THAT MAY BE APPLIED TO THE ELECTRODE. TABLE 2 COMPOSITION OF CARBON STEEL WELD METAL MADE WITH COMPOSITE ELECTRODE

ELECTRODE AWS COMPOSITION, WT%(A)(B)(C) CLASSIFICATION C Mn Si S P Cu EC1 0.15 1.80 0.90 0.035 0.035 0.35 Source: Ref 12

(A) WELD METAL SHALL BE ANALYZED FOR THE SPECIFIC ELEMENTS FOR WHICH VALUES ARE SHOWN IN THIS TABLE. IF THE PRESENCE OF OTHER ELEMENTS IS INDICATED, IN THE COURSE OF THIS WORK, THE AMOUNT OF THOSE ELEMENTS SHALL BE DETERMINED TO ENSURE THAT THEIR TOTAL (EXCLUDING IRON) DOES NOT EXCEED 0.50%. (B) SINGLE VALUES ARE MAXIMUM. (C) A LOW-DILUTION AREA OF A GROOVE WELD OR A FRACTURED TENSION TEST SPECIMEN AS DESCRIBED IN AWS A5.17 (SEE REF 10) MAY BE SUBSTITUTED FOR THE WELD PAD, AND SHALL MEET THE ABOVE REQUIREMENTS. IN CASE OF DISPUTE, THE WELD PAD SHALL BE THE REFEREE METHOD.

FIG. 3 CLASSIFICATION SYSTEM APPLICATIONS. SOURCE: REF 12

FOR

CARBON

STEEL

ELECTRODES

AND

FLUXES

USED

IN

SAW

F7A6-EM12K is a complete designation. It refers to a flux that will produce weld metal that, in the as-welded condition, will have a tensile strength no lower than 485 MPa (70 ksi) and Charpy V-notch impact strength of at least 27 J (20 ft · lbf) at -51 °C (-60 °F) when produced with an EM12K electrode under the conditions called for in this specification. F7A4-EC1 is a complete designation for a flux when the trade name of the electrode used in classification is indicated as

well. It refers to a flux that will produce weld metal with that electrode, which in the as-welded condition, will have a tensile strength no lower than 485 MPa (70 ksi) and Charpy V-notch energy of at least 27 J (20 ft · lbf) at -40 °C (-40 °F) under the conditions called for in this specification.

Low-Alloy Steels. The requirements for low-alloy steel welding electrodes are detailed in Ref 13. The classification

system for flux-electrode combinations of low-alloy steel is shown in Fig. 4. Table 3 shows the AWS electrode classification as determined by the composition of the solid electrode composition, while Table 4 is the flux-electrode classification requirement for the weld deposit composition of both solid and composite flux-electrode combinations. Interpretation of the classification system code is illustrated in the flux-electrode combination examples that follow.

TABLE 3 COMPOSITION OF LOW-ALLOY STEEL SOLID ELECTRODES ELECTRODE AWS CLASSIFICATI ON

UNS NO.

EL12(E)

K0101 2

EM12K(E)

K0111 3

EA1

K1122 2

EA2

K1122 3

EA3

K1142 3

EA3K

K2145 1

EA4

K1142 4

EB1

K1104 3

EB2

K1117 2

EB2H

K2301 6

EB3

K3111 5

EB5

K1218 7

EB6(F)

S5028 0

EB6H

S5018 0

COMPOSITION, WT%(A)(B)(C) C Mn Si S

0.04 0.14 0.05 0.15 0.07 0.17 0.07 0.17 0.07 0.17 0.07 0.12 0.07 0.17 0.10

0.07 0.15 0.28 0.33 0.05 0.15 0.18 0.23 0.10

0.25 0.40

0.25 0.60 0.80 1.25 0.65 1.00 0.95 1.35 1.65 2.20 1.60 2.10 1.20 1.70 0.40 0.80 0.45 1.00 0.45 0.65 0.40 0.80 0.40 0.70 0.35 0.70 0.75 1.00

P

Cr

Ni

Mo

Cu(d)

V

Al

Ti

Zr

0.10

0.03 0

0.03 0

...

...

...

0.35

...

...

...

...

0.10 0.35 0.20

0.03 0

0.03 0

...

...

...

0.35

...

...

...

...

0.03 0

0.02 5

...

...

0.35

...

...

...

...

0.20

0.03 0

0.02 5

...

...

0.35

...

...

...

...

0.20

0.03 0

0.02 5

...

...

0.35

...

...

...

...

0.50 0.80 0.20

0.02 5

0.02 5

...

...

0.35

...

...

...

...

0.03 0

0.02 5

...

...

0.35

...

...

...

...

0.05 0.30 0.05 0.30 0.55 0.75 0.05 0.30 0.40 0.60 0.05 0.50 0.25 0.50

0.02 5

0.02 5

...

0.35

...

...

...

...

0.03 0

0.02 5

0.35

...

...

...

...

0.01 5

0.01 5

0.30

...

...

0.02 5

0.35

0.20 0.30 ...

...

0.02 5

...

...

...

0.02 5

0.02 5

0.30

...

...

...

...

0.02 5

0.02 5

0.35

...

...

...

...

0.03 0

0.02 5

0.40 0.75 1.00 1.75 1.00 1.50 2.25 3.00 0.45 0.65 4.50 6.50 4.80 6.00

0.45 0.65 0.45 0.65 0.45 0.65 0.40 0.60 0.45 0.65 0.45 0.65 0.45 0.65 0.40 0.65 0.90 1.10 0.90 1.20 0.45 0.70 0.45 0.65

0.35

...

...

...

...

...

...

...

...

...

...

ELECTRODE AWS CLASSIFICATI ON

UNS NO.

COMPOSITION, WT%(A)(B)(C) C Mn Si S

EB8(F)

S5048 0

0.10

ENI1

K1104 0

0.12

ENI2

K2101 0

ENI3

K3131 0

ENI4

K1148 5

ENIIK

K1105 8

EF1

K1116 0

EF2

K2145 0

EF3

K2148 5

EF4

K1204 8

EF5

K4137 0

EF6

K2113 5

EM2(G)

K1088 2

EM3(G)

K2101 5

EM4(G)

K2103 0

EW

K1124 5

EG

...

0.30 0.65

P

Cr

Ni

Mo

Cu(d)

V

Al

Ti

Zr

8.00 10.5 0 0.15

...

0.80 1.20

0.35

...

...

...

...

0.75 1.25 2.10 2.90 3.10 3.80 1.60 2.10 0.75 1.25 0.95 1.60 0.40 0.80 0.70 1.10 0.40 0.80 2.30 2.80 1.75 2.25 1.40 2.10 1.90 2.60 2.00 2.80 0.40 0.80

0.30

0.35

...

...

...

...

...

0.35

...

...

...

...

...

0.35

...

...

...

...

0.10 0.30 ...

0.35

...

...

...

...

0.35

...

...

...

...

0.25 0.55 0.40 0.65 0.40 0.65 0.15 0.30 0.45 0.65 0.40 0.65 0.25 0.55 0.25 0.65 0.30 0.65 ...

0.35

...

...

...

...

0.35

...

...

...

...

0.35

...

...

...

...

0.35

...

...

...

...

0.50

...

...

...

...

0.35

...

...

...

...

0.25

0.05

0.1 0

0.1 0

0.1 0

0.25

0.04

0.1 0

0.1 0

0.1 0

0.25

0.03

0.1 0

0.1 0

0.1 0

0.30 0.80

...

...

...

...

0.05 0.50

0.03 0

0.04 0

0.75 0.05 1.25 0.30 0.12 0.75 0.05 1.25 0.30 0.13 0.60 0.05 1.20 0.30 0.12 0.60 0.10 0.19 1.00 0.30 0.12 0.80 0.40 1.40 0.80 0.07 0.90 0.15 0.15 1.70 0.35 0.10 1.70 0.20 0.18 2.40 0.10 1.70 0.30 0.18 2.40 0.16 0.60 0.15 0.23 0.90 0.35 0.10 1.70 0.20 0.17 2.20 0.07 1.45 0.10 0.15 1.90 0.30 0.10 1.25 0.20 1.80 0.60 0.10 1.40 0.20 1.80 0.60 0.10 1.40 0.20 1.80 0.60 0.12 0.35 0.20 0.65 0.35 NOT SPECIFIED

0.02 0

0.02 0

0.02 0

0.02 0

...

0.02 0

0.02 0

0.15

0.02 0

0.01 5

...

0.02 0

0.02 0

...

0.02 5

0.02 5

...

0.02 5

0.02 5

...

0.02 5

0.02 5

...

0.03 5

0.02 5

0.40 -.60

0.01 0

0.01 0

0.01 5

0.01 5

0.01 0

0.01 0

0.25 0.50 0.20 0.55 0.30

0.01 0

0.01 0

0.55

0.01 0

0.01 0

0.60

0.04 0

0.03 0

0.50 0.80

Source: Ref 13

(A) THE FILLER METAL SHALL BE ANALYZED FOR THE SPECIFIC ELEMENTS FOR WHICH VALUES ARE SHOWN IN THIS TABLE. IF THE PRESENCE OF OTHER ELEMENTS IS INDICATED, IN THE COURSE OF THIS WORK, THE AMOUNT OF THOSE ELEMENTS SHALL BE DETERMINED TO ENSURE THAT THEIR TOTAL (EXCLUDING IRON) DOES NOT EXCEED 0.50 WT%. (B) SINGLE VALUES ARE MAXIMUM. (C) THE LETTER "N" AS A SUFFIX TO A CLASSIFICATION INDICATES THAT THE ELECTRODE IS INTENDED FOR WELDS IN THE CORE BELT REGION OF NUCLEAR REACTOR VESSELS. THIS SUFFIX CHANGES THE LIMITS ON THE PHOSPHORUS, VANADIUM, AND COPPER AS FOLLOWS: P = 0.012% MAXIMUM: V = 0.05% MAXIMUM; CU = 0.08% MAXIMUM. "N" ELECTRODES SHALL NOT BE COATED WITH COPPER OR ANY MATERIAL CONTAINING COPPER. THE "EF5" AND "EW" ELECTRODES SHALL NOT BE DESIGNATED AS "N" ELECTRODES. (D) THE COPPER LIMIT INCLUDES ANY COPPER COATING THAT MAY BE APPLIED TO THE ELECTRODE. (E) THE EL12 AND EM12K CLASSIFICATIONS ARE IDENTICAL TO THOSE SAME CLASSIFICATIONS IN ANSI/AWS A5.17-89 (REF. 10). THEY ARE INCLUDED IN THIS SPECIFICATION BECAUSE THEY ARE SOMETIMES USED WITH AN ALLOY FLUX TO DEPOSIT SOME OF THE WELD METALS CLASSIFIED IN TABLE 4. (F) THE EB6 AND EB8 CLASSIFICATIONS ARE SIMILAR, BUT NOT IDENTICAL, TO THE ER502 AND ER505 CLASSIFICATIONS, RESPECTIVELY, IN ANSI/AWS A5.9-80, "SPECIFICATION FOR CORROSION-RESISTING CHROMIUM AND CHROMIUM-NICKEL STEEL BARE AND COMPOSITE METAL CORED AND STRANDED ARC WELDING ELECTRODES AND WELDING RODS." THESE CLASSIFICATIONS WILL BE DROPPED FROM THE NEXT REVISION OF A5.9 (SEE REF 14). (G) THE COMPOSITION RANGES OF CLASSIFICATIONS WITH THE "EM" PREFIX ARE INTENDED TO CONFORM TO THE RANGES FOR SIMILAR ELECTRODES IN THE MILITARY SPECIFICATIONS. TABLE 4 COMPOSITION OF LOW-ALLOY STEEL WELD METAL (BOTH SOLID FLUX-ELECTRODE AND COMPOSITE FLUX-ELECTRODE COMBINATIONS) AWS WELD METAL CLASSIFICATI ON(E)

UNS NO.

COMPOSITION, WT%(A)(B)(C)(D)

SOLID

COMPOSI TE

C

Mn

Si

S

A1

K112 22

W17041

0.1 2

1.0 0

0.8 0

A2

K112 23

W17042

0.1 2

1.4 0

A3

K114 23

W17043

0.1 5

A4

K114 24

W17044

0.1 5

P

Cr

Ni

Mo

Cu

V

Ti

Z r

0.04 0.03 . . . 0 0

...

0.3 5

...

...

. . .

0.8 0

0.04 0.03 . . . 0 0

...

0.3 5

...

...

. . .

2.1 0

0.8 0

0.04 0.03 . . . 0 0

...

0.3 5

...

...

. . .

1.6 0

0.8 0

0.04 0.03 . . . 0 0

...

0.4 00.6 5 0.4 00.6 5 0.4 00.6 5 0.4 00.6 5

0.3 5

...

...

. . .

B1

K110 43

W51040

0.1 2

1.6 0

0.8 0

0.04 0.03 0.40 . . . 0 0 0.65

B2

K111 72

W52040

0.1 5

1.2 0

0.8 0

0.04 0.03 1.00 . . . 0 0 1.50

B2H

K230 16

W52240

1.2 0

0.8 0

0.04 0.03 1.00 . . . 0 0 1.50

B3

K311 15

W53040

0.1 00.2 5 0.1 5

1.2 0

0.8 0

0.04 0.03 2.00 . . . 0 0 2.50

B4

...

W53346

0.1 2

1.2 0

0.8 0

0.04 0.03 1.75 . . . 0 0 2.25

B5

K121 87

W51348

0.1 8

1.2 0

0.8 0

0.04 0.03 0.40 . . . 0 0 0.65

B6

S5028 W50240 0

0.1 2

1.2 0

0.8 0

0.04 0.03 4.50 . . . 0 0 6.00

B6H

S5018 W50140 0

1.2 0

0.8 0

0.04 0.03 4.50 . . . 0 0 6.00

B8

S5018 W50440 0

0.1 00.2 5 0.1 2

1.2 0

0.8 0

NIL(F)

K110 40

W21048

0.1 2

1.6 0

0.8 0

NI2(F)

K210 10

W22040

0.1 2

1.6 0

0.8 0

NI3

K313 10

W23040

0.1 2

1.6 0

0.8 0

NI4

K114 85

W21250

0.1 4

1.6 0

0.8 0

F1

K111

W21150

0.1

0.7

0.8

0.04 0.03 8.00 . . . 0 0 10.0 0 0.03 0.03 0.15 0.7 0 0 51.1 0 0.03 0.03 . . . 2.0 0 0 02.9 0 0.03 0.03 0.15 2.8 0 0 03.8 0 0.03 0.03 . . . 1.4 0 0 02.1 0 0.04 0.03 0.15 0.9

0.3 5

...

...

. . .

0.3 5

...

...

. . .

0.3 5

0.3 0

...

. . .

0.3 5

...

...

. . .

0.3 5

...

...

. . .

0.3 0

...

...

. . .

0.3 5

...

...

. . .

0.3 5

...

...

. . .

0.3 5

...

...

. . .

0.3 5

...

0.05(

. . .

...

0.3 5

...

...

. . .

...

0.3 5

...

...

. . .

0.3 5

0.3 5

...

...

. . .

0.5

0.3

...

...

.

0.4 00.6 5 0.4 00.6 5 0.4 00.6 5 0.9 01.2 0 0.4 00.6 5 0.9 01.2 0 0.4 00.6 5 0.4 00.6 5 0.8 01.2 0 0.3 5

G)

60

2

F2

K214 50

W20240

0.1 7

F3

K214 85

W21140

0.1 7

F4

K120 48

W20440

0.1 7

F5

K413 70

W22640

0.1 7

F6

K211 35

W21040

0.1 4

M1

...

W21240

0.1 0

M2

M3

M4

...

...

...

W21340

W22240

W22440

W

...

W21040

G

NOT SPECIFIED

0.1 0

0.1 0

0.1 0

0.1 2

01.5 0 1.2 52.2 5 1.2 52.2 5 1.6 0

0

0

0

0.8 0

0.04 0.03 . . . 0 0

0.8 0

0.04 0.03 . . . 0 0

0.8 0

0.04 0.03 0.60 0 0

1.2 01.8 0 0.8 01.8 5 0.6 01.6 0 0.9 01.8 0 0.9 01.8 0 1.3 02.2 5 0.5 01.6 0

0.8 0

0.03 0.03 0.65 0 0

0.8 0

0.03 0.03 0.65 0 0

0.8 0

0.04 0.03 0.15 0 0

0.8 0

0.04 0.03 0.35 0 0

0.8 0

0.03 0.03 0.65 0 0

0.8 0

0.03 0.03 0.80 0 0

0.8 0

0.04 0.03 0.45 0 0 0.70

01.7 0 0.4 00.8 0 0.7 01.1 0 0.4 00.8 0 2.0 02.8 0 1.5 02.2 5 1.2 52.0 0 1.4 02.1 0 1.8 02.6 0 2.0 02.8 0 0.4 00.8 0

5

5

. .

0.4 00.6 5 0.4 00.6 5 0.2 5

0.3 5

...

...

. . .

0.3 5

...

...

. . .

0.3 5

...

0.03(

. . .

0.3 00.8 0 0.6 0

0.5 0

...

...

. . .

0.4 0

...

...

. . .

0.3 5

0.3 0

...

0.03(

. . .

0.2 50.6 5 0.2 00.7 0 0.3 00.8 0 ...

0.3 0

...

0.3 0

...

0.3 0

...

0.3 00.7 5

...

G)

G)

0.03( G)

0.03( G)

. . . . . .

0.03( G)

...

. . .

Source: Ref 13

(A) THE WELD METAL SHALL BE ANALYZED FOR THE SPECIFIC ELEMENTS FOR WHICH VALUES ARE SHOWN IN THIS TABLE. IF THE PRESENCE OF OTHER ELEMENTS IS INDICATED, IN THE COURSE OF THIS WORK, THE AMOUNT OF THOSE ELEMENTS SHALL BE DETERMINED TO ENSURE THAT THEIR TOTAL (EXCLUDING IRON) DOES NOT EXCEED 0.50%. (B) SINGLE VALUES ARE MAXIMUM.

(C) THE LETTER "N" AS A SUFFIX TO CLASSIFICATION INDICATES THAT THE WELD METAL IS INTENDED FOR WELDS IN THE CORE BELT REGION OF NUCLEAR REACTOR VESSELS. THIS SUFFIX CHANGES THE LIMITS ON PHOSPHORUS, VANADIUM, AND COPPER AS FOLLOWS: P = 0.012% MAXIMUM; V = 0.05% MAXIMUM; CU = 0.08% MAXIMUM. "N" ELECRODES SHALL NOT BE COATED WITH COPPER OR ANY MATERIAL CONTAINING COPPER. THE F5 AND W CLASSIFICATIONS SHALL NOT BE DESIGNATED AS "N" WELD METALS. (D) A LOW-DILUTION AREA OF THE GROOVE WELD OR THE REDUCED SECTION OF THE FRACTURED TENSION TEST SPECIMEN MAY BE SUBSTITUTED FOR THE WELD PAD, AND SHALL MEET THE ABOVE REQUIREMENTS. IN CASE OF DISPUTE, THE WELD PAD SHALL BE THE REFEREE METHOD. (E) THE ELECRODE DESIGNATION FOR COMPOSITE ELECTRODES IS OBTAINED BY PLACING AN "EC" BEFORE THE APPROPRIATE WELD METAL CLASSIFICATION. (F) MANGANESE IN NI1 AND NI2 CLASSIFICATIONS MAY BE 1.80% MAXIMUM WHEN CARBON IS RESTRICTED TO 0.10% MAXIMUM. (G) A SINGLE VALUE SPANNING THE COLUMNS FOR VANADIUM, TITANIUM, AND ZIRCONIUM IS THE TOTAL OF THESE ELEMENTS.

FIG. 4 CLASSIFICATION SYSTEM FOR LOW-ALLOY STEEL ELECTRODES AND FLUXES USED IN SAW APPLICATIONS

F9PO-EB3-B3 is a complete designation for a flux-electrode combination. It refers to a flux that will produce weld metal

that, in the postweld heat-treated condition, will have a tensile strength of 620 to 760 MPa (90 to 110 ksi) and Charpy Vnotch impact strength of at least 27 J (20 ft · lbf) at - 18 °C (0 °F) when made with an EB3 electrode under the conditions called for in this specification. The composition of the weld metal will be B3 (Table 4). F9A2-ECM1-M1 is a complete designation for a flux when the trade name of the composite electrode used in the

classification tests is indicated as well. The designation refers to a flux that will produce weld metal that, with the M1

electrode, in the as-welded condition, will have a tensile strength of 620 to 760 MPa (90 to 110 ksi) and Charpy V-notch impact strength of at least 27 J (20 ft · lbf) at -29 °C (-20 °F) under the conditions called for in this specification. The composition of the weld metal will be M1 (Table 4). Stainless and Nickel-Base Alloys. For welding with corrosion-resistant chromium and chromium-nickel bare and

composite metal cored and stranded electrodes, the composition of the filler metal is classified according to AWS A5.9 (Table 5). There is no AWS classification of fluxes or of flux-electrode combinations for stainless steel SAW.

TABLE 5 COMPOSITION OF CORROSION-RESISTANT CHROMIUM AND CHROMIUM-NICKEL BARE AND COMPOSITE METAL CORED AND STRANDED ELECTRODES AWS ELECTRODE CLASSIFICATION

COMPOSITION, WT%(A)(B) C Cr Ni

ER209(D)

0.05

ER218

0.10

ER219

0.05

ER240

0.05

ER307

0.040.14 0.08

ER308(C) ER308H ER308L(C)

0.040.08 0.03

ER308MO

0.08

ER308MOL

0.04

ER309(D)

0.12

ER309L

0.03

ER310 ER312

0.080.15 0.15

ER316(C)

0.08

ER316H ER316L(D)

0.040.08 0.03

ER317

0.08

ER317L

0.03

ER318

0.08

20.524.0 16.018.0 19.021.5 17.019.0 19.522.0 19.522.0 19.522.0 19.522.0 18.021.0 18.021.0 23.025.0 23.025.0 25.028.0 28.032.0 18.020.0 18.020.0 18.020.0 18.520.5 18.520.5 18.0-

9.512.0 8.09.0 5.57.0 4.06.0 8.010.7 9.011.0 9.011.0 9.011.0 9.012.0 9.012.0 12.014.0 12.014.0 20.022.5 8.010.5 11.014.0 11.014.0 11.014.0 13.015.0 13.015.0 11.0-

Mo

(nb + ta)

1.53.0 0.75

...

0.75 0.75 0.51.5 0.75 0.75 0.75 2.03.0 2.03.0 0.75 0.75 0.75 0.75 2.03.0 2.03.0 2.03.0 3.04.0 3.04.0 2.0-

Mn

4.07.0 ... 7.09.0 ... 8.010.0 ... 10.513.5 ... 3.34.75 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 ... 1.02.5 8 × C 1.0-

Si

P

S

N

Cu

0.90

0.03

0.03

0.75

3.54.5 1.00

0.03

0.03

0.03

0.03

1.00

0.03

0.03

0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.30065 0.300.65 0.300.65 0.300.65 0.300.65 0.300.65 0.30-

0.03

0.03

0.100.30 0.080.18 0.100.30 0.100.20 ...

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.75 0.75 0.75 0.75

20.0

14.0

3.0

2.5

0.65

2.5

0.60

0.03

...

3.0-4.0

1.52.0

0.15

0.015 0.020 . . .

3.0-4.0

1.02.5 ... 1.02.5 10 × C 1.02.5 MIN TO 1.0 MAX 1.0-1.4 1.02.5 ... 0.6

0.300.65 0.300.65 0.300.65

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.03

0.03

...

0.75

0.300.65 0.50

0.03

0.03

...

0.75

0.03

0.03

...

0.75

MIN TO 1.0 MAX 8 × C MIN TO 1.0 MAX 8 × C MIN TO 0.40 MAX ...

0.03

ER320

0.07

19.021.0

32.036.0

2.03.0

ER320LR(D)

0.025

19.021.0

32.036.0

2.03.0

ER321(E)

0.08 0.180.25 0.08

9.010.5 34.037.0 9.011.0

0.75

ER330

18.520.5 15.017.0 19.021.5

19.021.5 11.513.5 11.012.5 12.014.0 15.517.0 4.66.0 8.010.5 16.016.75 25.027.5 14.516.5

8.09.5 0.6

0.350.65 0.75

4.05.0 0.6

0.4.0.7 0.75

...

0.6

0.50

0.03

0.03

...

0.75

...

0.6

0.50

0.03

0.03

...

0.75

0.6

0.75

...

0.6

0.50

0.03

0.03

...

0.75

0.6

0.450.65 0.81.2 0.75

...

0.6

0.50

0.03

0.03

...

0.75

...

0.6

0.50

0.04

0.03

...

0.75

0.150.30 ...

0.250.75 0.40

0.75

0.04

0.03

...

0.40

0.02

0.02

0.015

3.254.00 0.20(G)

...

1.02.5

0.300.65

0.03

0.03

...

0.75

ER347(C)

ER349(F) ER410

0.070.13 0.12

ER410NIMO

0.06

ER420 ER430

0.250.40 0.10

ER502

0.10

ER505

0.10

ER630

0.05

ER26-1

0.01

ER16-8-2

0.10

0.5 4.55.0 (G)

7.59.5

0.75 0.75

0.751.50 1.02.0

Source: Ref 14

(A) ANALYSIS SHALL BE MADE FOR THE ELEMENTS FOR WHICH SPECIFIC VALUES ARE SHOWN IN THIS TABLE. IF, HOWEVER, THE PRESENCE OF OTHER ELEMENTS IS INDICATED IN THE COURSE OF ROUTINE ANALYSIS, FURTHER ANALYSIS SHALL BE MADE TO DETERMINE THAT THE TOTAL OF THESE OTHER ELEMENTS, EXCEPT IRON, IS NOT PRESENT IN EXCESS OF 0.50 %. (B) SINGLE VALUES SHOWN ARE MAXIMUM PERCENTAGES EXCEPT WHERE OTHERWISE SPECIFIED. (C) THESE GRADES ARE AVAILABLE IN HIGH-SILICON CLASSIFICATIONS THAT SHALL HAVE THE SAME CHEMICAL COMPOSITION REQUIREMENTS AS GIVEN BELOW WITH THE EXCEPTION THAT THE SILICON CONTENT SHALL BE 0.65 TO 1.00%. THESE HIGH-SILICON CLASSIFICATIONS SHALL BE DESIGNATED BY THE ADDITION `SI' TO THE STANDARD

CLASSIFICATION DESIGNATIONS INDICATED BELOW. THE FABRICATOR SHOULD CONSIDER CAREFULLY THE USE OF HIGH-SILICON FILLER METALS IN HIGHLY RESTRAINED FULLY AUSTENITIC WELDS. A DISCUSSION OF THE PROBLEM IS PRESENTED IN REF 14. (D) CARBON SHALL BE REPORTED TO THE NEAREST 0.01% EXCEPT FOR THE CLASSIFICATION E320LR, FOR WHICH CARBON SHALL BE REPORTED TO THE NEAREST 0.005%. (E) TITANIUM: 9 × C MIN TO 1.0 MAX. (F) TITANIUM: 0.10 TO 0.30. TUNGSTEN IS 1.25 TO 1.75%. (G) NICKEL, MAX: 0.5 MINUS THE COPPER CONTENT. Like stainless steels, nickel and nickel alloys are classified according to the composition of the filler metal only (Ref 15). Classifications of filler metal are shown in Table 6. There is no AWS classification of fluxes or of flux-electrode combinations for nickel alloy SAW.

TABLE 6 COMPOSITION OF NICKEL AND NICKEL-ALLOY BARE ELECTRODES AND RODS

AWS CLASSIFICATIO N

UNS NO.

COMPOSITION, WT%(A)(B) C

MN

FE

P

S

SI

CU

N(C)

CO

AL

TI

CR

(NB + TA)

MO

V

W

OTHER ELEMENT S, TOTAL

ERNI-1

N0206 1

0.15

1.0

1.0

0.03

0.01 5

0.7 5

0.25

...

1.5

2.03.5

...

...

...

...

...

0.50

ERNICU-7

N0406 0

0.15

4.0

2.5

0.02

0.01 5

1.2 5

BA L

...

1.25

1.53.0

...

...

...

...

...

0.50

ERNICR-3

N0608 2

0.10

2.5- 3.0 3.5

0.03

0.01 5

0.5 0

0.50

93.0 MI N 62.0 69.0 67.0 MI N 70.0 MI N 67.0 MI N 38.0 46.0 50.0 55.0 BA L

(D)

...

0.75

2.0- . . . 3.0(E

...

...

0.50

...

...

0.50

ERNLCRFE-5

N0606 2

0.08

1.0

6.010.0

0.03

0.01 5

0.3 5

...

...

2.03.5

...

0.20

0.60 -1.2

18.0 22.0 14.0 17.0 14.0 17.0 19.5 23.5

...

2.5

0.20 0.80 ...

0.65 1.15 ...

17.0 21.0 1.0

BA L

0.20

...

...

6.08.0

0.50

BA L

2.5

...

...

4.06.0

...

0.50

BA

1.0

...

...

1.0

...

0.50

ERNLCRFE-6

N0709 2

0.08

2.0- 8.0 2.7

0.03

0.01 5

0.3 5

0.50

ERNIFECR-1

N0806 5

0.05

1.0

0.03

0.03

0.5 0

1.50 -3.0

ERNIFECR-2(F)

N0771 8

0.08

0.3 5

22.0 MI N BA L

0.01 5

0.01 5

0.3 5

0.30

ERNIMO-1

N1000 1

0.08

1.0

4.07.0

0.02 5

0.03

1.0

0.50

ERNIMO-2

N1000 3

0.04- 1.0 0.08

5.0

0.01 5

0.02

1.0

0.50

ERNIMO-3

N1000 4

0.12

1.0

4.07.0

0.04

0.03

1.0

ERNIMO-7

N1066

0.02

1.0

2.0

0.04

0.03

0.1

(D)

...

...

)

1.5- . . . 3.0(E )

...

...

...

...

0.50

...

2.53.5

...

...

0.50

4.75 5.50 ...

2.80 3.30 26.0 30.0 15.0 18.0 23.0 26.0 26.0

...

...

0.50

0.20 0.40 0.50

1.0

0.50

0.50

0.50

0.60

1.0

0.50

...

1.0

0.50

5 ERNICRMO-1

N0600 7

ERNICRMO-2

N0600 2

ERNICRMO-3

N0662 5

ERNICRMO-4

0 0.05

L

1.0- 18.0 2.0 21.0 0.05- 1.0 17.0 0.15 20.0 0.10 0.5 5.0 0

0.04

0.03

1.0

1.52.5

BA L

2.5

...

...

0.04

0.03

1.0

0.50

BA L

0.50 -2.5

...

...

0.02

0.01 5

0.5 0

0.50

...

0.40

0.40

N1027 6

0.02

1.0

4.07.0

0.04

0.03

0.0 8

0.50

58.0 MI N BA L

2.5

...

...

ERNICRMO-7

N0645 5

0.01 5

1.0

3.0

0.04

0.03

0.0 8

0.50

BA L

2.0

...

0.70

ERNICRMO-8

N0697 5

0.03

1.0

BA L

0.03

0.03

1.0

0.71.20

...

...

ERNICRMO-9

N0698 5

0.01 5

1.0

0.04

0.03

1.0

1.52.5

5.0

...

0.70 1.50 ...

ERNICRMO-10

N0602 2

0.01 5

0.5 0

18.0 21.0 2.06.0

47.0 52.0 BA L

0.02

0.01 0

0.0 8

0.50

BA L

2.5

...

...

ERNICRMO-11

N0603 0

0.03

1.5

0.04

0.02

0.8 0

1.02.4

BA L

5.0

...

...

ERNICRCOMO-1

N0661 7

0.05- 1.0 0.15

13.0 17.0 3.0

0.03

0.01 5

1.0

0.50

BA L

10.0 15.0

0.80 1.50

0.60

21.0 23.5 20.5 23.0 20.0 23.0 14.5 16.5 14.0 18.0 23.0 26.0 21.0 23.5 20.0 22.5 28.0 31.5 20.0 24.0

1.75 2.50 ...

30.0 5.57.5

...

1.0

0.50

8.010.0

...

0.20 -1.0

0.50

8.010.0

...

...

0.50

15.0 17.0 14.0 18.0 5.07.0

0.35

3.04.5

0.50

...

0.50

0.50

...

...

0.50

0.50

6.08.0

...

1.5

0.50

...

12.5 14.5 4.06.0

0.35

2.54.5

0.50

...

1.54.0

0.50

8.010.0

...

...

0.50

3.15 4.15 ...

...

...

0.30 1.50 ...

Source: Ref 15

(A) THE FILLER METAL SHALL BE ANALYZED FOR THE SPECIFIC ELEMENTS FOR WHICH VALUES AVE SHOWN IN THIS TABLE. IN THE COURSE OF THIS WORK, IF THE PRESENCE OF OTHER ELEMENTS IS INDICATED, THE AMOUNT OF THOSE ELEMENTS SHALL HE DETERMINED TO ENSURE THAT THEIR TOTAL DOES NOT EXCEED THE LIMIT SPECIFIED FOR "OTHER ELEMENTS, TOTAL" IN

THE LAST COLUMN OF THE TABLE. (B) SINGLE VALUES ARE MAXIMUM, EXCEPT WHERE OTHERWISE SPECIFIED. (C) INCLUDES INCIDENTAL COBALT. (D) COBALT: 0.12 MAXIMUM, WHEN SPECIFIED. (E) TANTALUM: 0.30 MAXIMUM, WHEN SPECIFIED. (F) BORON IS 0.006% MAXIMUM.

Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

Personnel Considerations Employee Training. Submerged arc welding is an automatic or semiautomatic process carried out primarily in the flat or

horizontal position with the arc obscured from view and does not require a high level of skill. The primary expenditures both in cost and time are consumed in designing jigs and fixtures to hold and position the workpiece so that SAW can be performed. While the actual welding process does not require a high degree of skill, the setup procedure requires skilled personnel with an aptitude for fabricating the fixtures and clamps necessary to position the workpiece in an acceptable position (usually flat). Health and Safety in the Workplace. The main safety concerns in welding are electric shock, fume inhalation, and

burns from both the arc ultraviolet source and the hot-metal infrared source (Ref 16). To reduce the risk of electrical shock, all equipment must be properly grounded and welding cables must be in good condition. The fume is nearly eliminated in SAW. However, certain vaporized elements (for example, chromium, cobalt, manganese, nickel, and vanadium) can be potentially dangerous. To prevent exposure to these elements, the welding area should have adequate ventilation. In SAW, the arcflash and its accompanying spatter are nearly eliminated because of the slag and the depth of the flux pile. However, eye protection should always be worn. Additional information is available in the article "Safe Practices" in this Volume.

Reference cited in this section

16. SAFETY IN WELDING AND CUTTING, AWS, 1988 Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

SAW Parameters While SAW is the most inexpensive and efficient process for making large, long, and repetitive welds, much time and energy are required to prepare the joint. Care must be taken to line up all joints to have a consistent gap in groove welds and to provide backing plates and flux dams to prevent spillage of flux and molten metal. Once all the pieces are clamped or tacked in place, welding procedures and specifications should be consulted before welding begins. Procedural Variations and Effect on Weld Bead Characteristics Procedural variations in SAW include current, voltage, electrical stickout (distance from last electrical contact to plate), travel speed, and flux depth. Variation in any of these parameters will affect the shape and penetration of the weld, as well as the integrity of the weld deposit. Weld Current. Because the welding current controls such parameters as deposition rate, penetration, and dilution, it is

the most important welding variable. An increase in welding current at a constant voltage will decrease the flux-to-wire ratio, while a decrease in current will increase the flux-to-wire ratio. The effect of current variation on weld bead profile is shown in Fig. 5. Welds made at excessively low current will tend to have little penetration and higher width-to-depth ratios. Welds made at an excessively high current will have deep penetration, high dilution, more shrinkage, and excess buildup. Low current will also produce a less stable arc than higher currents.

FIG. 5 EFFECT OF VARIATION IN WELDING CURRENT ON WELD BEAD PROFILE. (A) EXCESSIVELY LOW CURRENT. (B) EXCESSIVELY HIGH CURRENT. (C) RECOMMENDED CURRENT

The direction of current flow will also affect the weld bead profile. The current may be direct with the electrode positive (reverse polarity), electrode negative (straight polarity), or alternating. Reverse polarity is most commonly used. For a given set of welding conditions, reverse polarity will produce wider beads with more penetration at a lower deposition rate than straight polarity. Straight-polarity welding will contribute to narrower beads with less penetration and more buildup. Because straight polarity reduces base plate dilution, it is frequently used in surfacing applications. For the same welding current, the deposition rate with straight polarity is higher than with reverse polarity. Straight polarity is preferred for poor fitup. The bead shape, penetration, and deposition rate for alternating current fall between those of straight and reverse polarity. Alternating Current is used when welding current exceeds 1000 A and on multiple-wire applications to reduce arc blow and arc interaction. In SAW, the current density in the electrode also plays a role in bead shape and penetration. Smaller-diameter electrodes with a high current density will produce narrower beads with deeper penetration than larger-diameter electrodes. Largerdiameter electrodes are able to bridge larger root openings. In cases where a given current can be achieved with two different electrode diameters, the smaller electrode will produce the higher deposition rate. Weld Voltage. Like current, welding voltage will affect the bead shape and the weld deposit composition. Increasing the

arc voltage at a constant current will increase the flux-to-wire electrode ratio, while decreasing the voltage will reduce the flux-to-electrode ratio. The effect of the magnitude of arc voltage on bead shape is shown in Fig. 6. Increasing the arc voltage will produce a longer arc length and a correspondingly wider, flatter bead with less penetration. Higher voltage will increase flux consumption, which could then change deposit composition and properties. Slightly increasing the arc voltage will help the weld to bridge gaps when welding in grooves. Excessively high voltage will produce a hat-shaped concave weld, which has low resistance to cracking and a tendency to undercut. Lower voltages will shorten the arc length and increase penetration. Excessively low voltage will produce an unstable arc and a crowned bead, which has an uneven contour where it meets the plate.

FIG. 6 EFFECT OF VARIATION IN WELDING VOLTAGE AT CONSTANT CURRENT ON WELD BEAD PROFILE. (A) EXCESSIVELY LOW VOLTAGE. (B) EXCESSIVELY HIGH VOLTAGE

Electrical Stickout. In SAW, the current flowing in the electrode between the contact tube and the arc (electrode

extension) will cause some I2R heating, resulting in a voltage drop across that length of electrode. This resistance heating and subsequent voltage drop can be used to obtain higher deposition rates. Normal electrode extension for solid SAW wire is approximately 8 to 12 times the electrode diameter. As this length increases at a constant current, so does the resistance heating and the melt-off rate. To compensate for the voltage drop and the increase in wire-feed speed, the voltage must be increased to obtain a properly shaped bead. Extending the electrode 20 to 40 times the diameter can

increase deposition rates by more than 50% (Ref 17). Although higher deposition rates can be achieved by extending the electrode, keeping the electrode aligned with the joint becomes increasingly difficult as the extension increases. Travel Speed. Variations in travel speed at a set current and voltage also affect bead shape. As welding speed is

decreased, heat input per length of joint increases, and the penetration and bead width increase. The penetration will increase until molten metal begins to flow under the arc and interfere with heat flow at excessively slow speeds. Excessively high travel speeds will promote a crowned bead as well as the tendency for undercut and porosity. Flux layer depth is another variable that will alter the appearance, penetration, and quality of a submerged arc weld. If

the flux layer is too deep, a greater-than-normal amount of flux will be melted, resulting in weld beads that are narrower than normal. Some surface imperfections may also appear because gases may be trapped by the deep flux layer. If the flux layer is too shallow, the arc will flash through, and the bead will have a rough appearance or porosity due to lack of shielding from the atmosphere. The correct depth of flux is just enough to prevent flash through. This will allow welding gases to escape while providing adequate protection. Sources of Defects in SAW The fact that SAW is a high heat input process under a protective blanket of flux greatly decreases the chance of weld defects. However, defects such as lack of fusion, slag entrapment, solidification cracking, hydrogen cracking, or porosity occasionally occur. Insufficient Fusion and Slag Entrapment. Lack-of-fusion defects and slag entrapment are most commonly caused by improper bead placement or procedure. Improper placement can cause the weld metal to roll over and trap slag underneath, or if the weld bead is placed away from the edge to be joined, the liquid metal may not fuse to the base material. A crown-shaped bead caused by low welding voltage may also contribute to slag entrapment and lack of fusion by not allowing the liquid metal to spread out evenly. Solidification cracking of SAW along the center of the bead is usually due to bead shape, joint design, or incorrect choice of welding consumables. A convex bead shape with a bead width-to-depth ratio greater than one will decrease solidification cracking tendencies. If weld penetration is too deep, the shrinkage stresses may cause centerline cracking. Joint design may also contribute to excessive shrinkage stresses, again increasing the risk of solidification cracking. Because cracking is related to stresses in the weld, high-strength materials will have a greater tendency to crack. Therefore, special care must be taken to generate proper bead shape, preheat temperatures, and interpass temperatures, in addition to correct electrode and flux combinations, when welding these materials. Hydrogen Cracking. Unlike solidification cracking, which appears immediately after welding, hydrogen cracking is a

delayed process and may occur from several hours to several days after welding has been completed. To minimize hydrogen cracking, all possible sources of hydrogen (for example, water, oil, grease, and dirt) present in the flux, electrode, or joint should be eliminated. The flux, electrode, and plate should be clean and dry. To prevent moisture pickup, fluxes and electrodes should be stored in moisture-resistant containers in dry areas. If a flux or electrode becomes contaminated with moisture, it should be dried according to manufacturer recommendations. Selection of the consumable, especially when welding high-strength steels that are more susceptible to hydrogen cracking, can also play a role in hydrogen cracking. Special consumables have recently been developed that produce weld deposits that are very low in diffusible hydrogen. To further reduce hydrogen-related cracking, the joint to be welded should be preheated. Because hydrogen is fairly mobile in steel at temperatures above 95 °C (200 °F), the recommended preheat temperatures should be followed to allow most of the hydrogen to escape and to reduce the risk of hydrogen damage (Ref 18). In thick weldments, maintaining preheat for several hours after welding is completed will also reduce the risk of hydrogen cracking. Porosity caused by trapped gas is uncommon in SAW because of the protection provided by the flux. When porosity

does occur, it may be in the form of internal porosity or as depressions on the weld bead surface. The gas bubbles that cause porosity originate either from a lack of protection from the atmosphere or from contaminants such as water, oil, grease, and dirt. To reduce porosity in SAW, the weld should have sufficient flux coverage, and all water, grease, and dirt should be removed from the plate, electrode, and flux. Another cause of porosity in SAW is excessive travel speed. Travel

at excessively high speeds will not allow the gas bubbles to escape from the weld, and the bubbles may become trapped in the weld metal at the slag-to-metal interface.

References cited in this section

17. HOW TO MAKE SINGLE ELECTRODE SUBMERGED ARC WELDS, PUBLICATION S604, THE LINCOLN ELECTRIC CO., 1991, P 11 18. F.R. COE, WELD. WORLD, VOL 14 (NO. 1/2), 1976, P 1-7 Submerged Arc Welding Jonathan S. Ogborn, The Lincoln Electric Company

References

1. R.L. O'BRIEN, WELDING HANDBOOK, VOL 11, AWS, 1991, P 191-232 2. D.L. OLSON ET AL., SUBMERGED ARC WELDING, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 114-152 3. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 6.3-1 TO 6.3-24 4. "SPECIFICATION FOR CARBON STEEL ELECTRODES AND FLUXES FOR SUBMERGED ARC WELDING," A5.17-89, AWS, 1989, P 18 5. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 3.2-3, 6.3-21 TO 6.3-22, 6.3-49 TO 6.3-58 6. THE PROCEDURE HANDBOOK OF ARC WELDING, THE LINCOLN ELECTRIC CO., 1973, P 6.3-26 TO 6.3-73 7. R.S. BROWN ET AL., ARC WELDING OF STAINLESS STEELS, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 327-329 8. K.C. ANTONY ET AL., HARDFACING, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 784 9. D.R. THOMAS ET AL., WELD OVERLAYS, VOL 6, 9TH ED., METALS HANDBOOK, AMERICAN SOCIETY FOR METALS, 1983, P 804-819 10. "SPECIFICATION FOR CARBON STEEL ELECTRODES AND FLUXES FOR SUBMERGED ARC WELDING," A5.17-89, AWS, 1989, P 17-18 11. S.S. TULIANI, T. BONISZEWSKI, AND N.F. EATON, NOTCH TOUGHNESS OF COMMERCIAL SUBMERGED ARC WELD METAL, WELD. MET. FABR., VOL 8, 1969, P 327-339 12. "SPECIFICATION FOR CARBON STEEL ELECTRODES AND FLUXES FOR SUBMERGED ARC WELDING," A5.17-89, AWS, 1989, P 2-3 13. "SPECIFICATION FOR LOW ALLOY STEEL ELECTRODES AND FLUXES FOR SUBMERGED ARC WELDING," A5.23-90, AWS, 1990, P 2-6 14. "SPECIFICATION FOR CORROSION RESISTING CHROMIUM AND CHROMIUM-NICKEL STEEL BARE AND COMPOSITE METAL CORED AND STRANDED WELDING ELECTRODES AND WELDING RODS," A5.9-81, AWS, 1981, P 3, 4 15. "SPECIFICATION FOR NICKEL AND NICKEL ALLOY BARE WELDING ELECTRODES AND RODS," A5.14-89, AWS, 1989, P 2, 3 16. SAFETY IN WELDING AND CUTTING, AWS, 1988 17. HOW TO MAKE SINGLE ELECTRODE SUBMERGED ARC WELDS, PUBLICATION S604, THE LINCOLN ELECTRIC CO., 1991, P 11 18. F.R. COE, WELD. WORLD, VOL 14 (NO. 1/2), 1976, P 1-7

Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Introduction STUD ARC WELDING (SW), also known as arc stud welding, is a commonly used method for joining a metal stud, or fastener, to a metal workpiece. The process has been used as an alternative metal-fastening method since the 1940s. Millions of specially designed and manufactured metal studs are welded by this process every week in such diverse industries as construction, shipbuilding, automotive, and hard goods, as well as in miscellaneous industrial applications. This article will serve as a basic information source for those interested in accomplishing one-sided, no-hole attachment of metal fasteners. The SW process represents an alternative to other welding processes, and is also a substitute for other fastening procedures, such as drilling and tapping, bolting, and self-tapping screws. Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Process Overview Stud arc welding is similar to many other welding processes, including arc and percussion welding, in that the base (weld end) of a specifically designed stud is joined to a base material by heating both parts with an arc that is drawn between the two. Equipment that is unique to this process regulates the arc length and arc dwell time. After an arc is struck, the stud weld end and the workpiece surface are brought to the proper temperature for joining and, after a controlled period of time, the two heated surfaces are brought together under pressure, creating a metallurgical bond capable of developing the full strength of the stud. There are two basic types of stud arc welding, which are differentiated by the source of welding power. One type uses direct current (dc) power provided by a transformer/rectifier or a motor generator similar to that used in the shielded metal arc welding (SMAW) process. The second type uses power discharged from a capacitor storage bank. The process based on a dc power source is known as stud arc welding, whereas the process that utilizes capacitors is known as capacitor discharge stud welding (CDSW). Both the SW and CDSW processes overlap in some areas of application. Generally, the SW process is used in applications that require similar stud and workpiece metals, the workpiece thickness is greater in relation to the stud diameter, and an accommodation must be made for the stud flash (fillet). The term "flash" is preferred to the term "fillet," because the metal that forms the flash during the stud arc welding process is expelled, rather than added, as occurs with other welding processes. In contrast, the CDSW process is used extensively when welding to thin sheet metal, and is used frequently with dissimilar workpiece and stud alloys. It is also used in cases where marks on the opposite side of the workpiece must be avoided or minimized. With this process, the stud diameter is limited to smaller sizes. The CDSW process is more fully described in the article "Capacitor Discharge Stud Welding" in this Volume. The factors on which process selection should be based are fastener size, base-metal thickness, base-metal composition, and reverse-side marking requirements (Table 1).

TABLE 1 STUD-WELDING PROCESS SELECTION PARAMETERS

STUD SHAPE ROUND SQUARE

STUD ARC WELDING

CAPACITOR DISCHARGE STUD WELDING GAP AND DRAWN ARC METHOD CONTACT METHODS

A A

A A

A A

RECTANGULAR IRREGULAR STUD DIAMETER OR AREA

A A

A A

A A

1 1 TO IN.) 16 8 1 1 3.2 TO 6.4 MM ( TO IN.) 8 4 1 1 6.4 TO 12.7 MM ( TO IN.) 4 2 1 12.7 TO 25.4 MM ( TO 1 IN.) 2

D

A

A

C

A

A

A

B

B

A

D

D

UP TO 32.3 MM2 (0.05 IN.2) OVER 32.3 MM2 (0.05 IN.2) STUD METAL CARBON STEEL STAINLESS STEEL ALLOY STEEL ALUMINUM BRASS BASE METAL CARBON STEEL STAINLESS STEEL ALLOY STEEL ALUMINUM BRASS BASE-METAL THICKNESS UNDER 0.4 MM (0.015 IN.) 0.4 TO 1.6 MM (0.015 TO 0.062 IN.) 1.6 TO 3.2 MM (0.062 TO 0.125 IN.) OVER 3.2 MM (0.125 IN.) STRENGTH CRITERIA HEAT EFFECT ON EXPOSED SURFACES WELD FILLET CLEARANCE STRENGTH OF STUD GOVERNS STRENGTH OF BASE METAL GOVERNS

C A

A D

A D

A A B B C

A A C A A

A A C A D

A A B B C

A A A A A

A A C A D

D C B A

A A A A

B A A A

B B A A

A A A A

A A A A

1.6 TO 3.2 MM (

A. applicable without special procedures or equipment; B. applicable with special techniques or for specific applications that justify preliminary trials or testing to develop welding procedure and technique; C, limited application; D, not recommended-welding methods not developed at this time. Source: Ref 1 Equipment. As shown in Fig. 1, the basic equipment used for stud arc welding consists of a control system, which

regulates the arc time and controls gun movement; a fixed or portable stud-welding gun, which holds the stud in position during the welding process to create the proper arc length and joining pressure; and connecting cables, which must be connected to a separate source of dc power. The other items that are needed to weld the workpiece are the studs themselves and ceramic arc shields, or ferrules. The equipment used for stud welding is comparable, in terms of size, portability, and ease of operation, to the equipment used in the SMAW process. Although the initial cost of stud-welding systems varies with the method selected, stud size, and productivity requirements, it is generally competitive with other fastening methods on an in-place or finished part cost basis.

FIG. 1 STUD ARC WELDING CONTROL SYSTEM. THE CONTROL SYSTEM MUST BE CONNECTED TO A DC POWER SOURCE FOR WELDING. IN MOST APPLICATIONS, THE STUD (ELECTRODE) SHOULD BE NEGATIVE.

A stud-welding control system that has been integrated in a transformer/rectifier power source is shown in Fig. 2. This type of equipment, which is the most widely used, is called a power/control system. It can weld studs with diameters up to 28.5 mm (1

1 in.). Either a single gun or dual guns can be used. Although a light-duty control system can weigh 11.3 kg 8

(25 lb), the system shown in Fig. 2 weighs approximately 450 kg (1000 lb) and can be put on a wheeled cart for mobility.

FIG. 2 TYPICAL INTEGRATED POWER/CONTROL SYSTEM FOR STUD ARC WELDING

A typical stud-welding gun (Fig. 3) is usually available in at least two sizes: standard duty, for use with studs up to 16 mm (

5 in.) in diameter, and heavy duty, for use on larger-diameter studs. The weight of the gun can vary from approximately 8

1.8 kg (4 lb), for a standard-duty gun, to 6.8 kg (15 lb), for heavy-duty models.

FIG. 3 TYPICAL STANDARD-DUTY GUN FOR STUD ARC WELDING

Operation. The stud arc welding process utilizes the same principles as any other arc welding procedure. First, the stud,

which acts as an electrode, is inserted into a chuck on the end of the gun, surrounded by a ceramic ferrule, and positioned against the workpiece (Fig. 4a). Next, the gun trigger is depressed, which starts the automatic weld cycle by energizing a solenoid coil within the gun body that lifts the stud off the work and draws an arc. The arc melts the end of the stud and a portion of the workpiece (Fig. 4b). After a preset arc time (set on the control unit), the welding current is shut off and the solenoid is de-energized (Fig. 4c). A mainspring in the gun forces the stud into the molten pool of metal, producing a fullstrength weld, which is shown in Fig. 4(d) after the gun has been lifted off the stud and the ceramic ferrule removed. The result is a full-penetration, full-strength stud-to-workpiece weld, as shown in Fig. 5. This weld is similar to that obtained with other types of arc-welding processes.

FIG. 4 STUD ARC WELDING PROCESS. (A) GUN IS PROPERLY POSITIONED. (B) TRIGGER IS DEPRESSED AND STUD IS LIFTED, CREATING AN ARC. (C) ARCING PERIOD IS COMPLETED AND STUD IS PLUNGED INTO MOLTEN POOL OF METAL ON BASE MATERIAL. (D) GUN IS WITHDRAWN FROM WELDED STUD AND FERRULE IS REMOVED.

FIG. 5 MACROSECTION OF LOW-CARBON STEEL STUD WELD

Studs and Ferrules. With the exception of special processes, the stud arc welding of steel, stainless steel, and aluminum

studs requires the use of a specifically designed weld stud and ceramic ferrule. Studs. A wide range of studs are made by stud-welding manufacturers for the SW process. They vary in size and weld-

base configuration, depending on the application requirements. The typical stud styles that are available include threaded, unthreaded, headed, and rectangular studs (Fig. 6). The steel stud diameters that are commonly welded using the SW process range from 2.7 to 25.4 mm (0.105 to 1.00 in.).

FIG. 6 COMMON STUD CONFIGURATIONS FOR STUD ARC WELDING

Full-strength welds result when using studs made from low-carbon steel that conforms to ASTM A-108 grades C-1010 through C-1020, although other steel grades can be used on a special-application qualified basis (Ref 2). Stainless steel studs are typically made from the austenitic series, including AISI 302, 302 HQ, 304, 305, 308, 309, 310, 316, 321, and 347, in both normal-carbon and low-carbon varieties. Stainless steel 303 is not an acceptable stud alloy. Aluminum studs are usually made from aluminum-magnesium alloys, including 5183, 5356, 5556, 5086, and 5456 (Ref 3). Steel and stainless steel arc welded studs that are larger than 6.4 mm (0.250 in.) in diameter use a welding flux to stabilize the welding arc and to deoxidize the weld area. The most commonly used flux is commercially pure aluminum, which is installed onto the weld end of the stud by thermal spraying, staking, or pressing a slug into a drilled hole in the end of the stud (Fig. 7). Studs that are 6.4 mm (0.250 in.) and under usually do not require fluxing.

FIG. 7 METHODS OF FLUX LOADING STUDS. (A) PRESSED IN SLUG. (B) STAKED ON WASHER. (C) COATING

Aluminum studs do not have to be flux loaded. Instead, the weld end of the stud has a conical or cylindrical tip that is shaped into the parent metal to help initiate the arc and control the arc length. To prevent the oxidation and embrittlement of the stud/plate weld zone during aluminum stud arc welding, the weld area is purged with an inert gas during the weld cycle. Argon gas is commonly used for this purpose, although helium also can be used. The typical gas flow rates for aluminum studs of various diameters are shown in Table 2.

TABLE 2 GAS FLOW RATES FOR ALUMINUM STUD ARC WELDING

STUD WELD BASE DIAMETER SHIELDING GAS FLOW(A) mm in. l/min ft3/h 1 6.4 7.1 15 7.9 9.5 11.1 12.7 (A)

4 5 16 3 8 7 16 1 2

7.1

15

9.4

20

9.4

20

9.4

20

SHIELDING GAS, 99.95% PURE ARGON

When stud arc welding low-carbon and stainless steel studs, the stud (electrode) should be negative. Aluminum stud arc welding should be set up so that the stud is positive. Obtaining a full-quality stud arc weld requires a sufficient amount of total energy input to the weld joint in order to produce melting and complete metallurgical bonding of the stud and workpiece. The energy input or weld current that is necessary depends on the stud diameter. Other parameters in the stud arc welding process are: arc voltage, arc time, and plunge. Arc voltage is a function of the arc length, which is set as "lift," or the length the stud is drawn away from the workpiece during weld-cycle initiation. Plunge is the length of stud that extends past the end of the ceramic ferrule and is available for melt off during the weld cycle. Studs are reduced in length during the weld cycle. Table 3 shows the typical reductions for steel and aluminum studs as a result of welding.

TABLE 3 TYPICAL STUD WELDING SETTINGS STUD BASE DIAMETE R mm in.

LENGTH REDUCTI ON mm

in.

DOWNHAND WELDING CURRE TIME LIFT NT, A s cycl m in. es m

STEEL AND STAINLESS STEEL 4.8 0.1 3. 0.12 300 0.1 87 2 5 5 6.4 0.2 3. 0.12 450 0.1 50 2 5 7 7.9 0.3 3. 0.12 500 0.2 12 2 5 5 9.5 0.3 3. 0.12 550 0.3 75 2 5 3 11. 0.4 3. 0.12 675 0.4 1 37 2 5 2 12. 0.5 4. 0.12 800 0.5 7 00 8 5 5 15. 0.6 4. 0.18 1200 0.6 9 25 8 7 7 19. 0.7 4. 0.18 1500 0.8 1 50 8 7 4 22. 0.8 4. 0.18 1700 1.0 2 75 8 7 0 25. 1.0 6. 0.25 1900 1.4 4 00 4 0 0 (A) ALUMINUM 4.8 0.1 3. 0.12 150 0.2 87 2 5 5 6.4 0.2 3. 0.12 200 0.4 50 2 5 0 7.9 0.3 3. 0.12 250 0.5 12 2 5 0 9.5 0.3 3. 0.12 325 0.6 75 2 5 5

8-9 1012 15 20 25 33 40 5055 6065 85

15 24 30 39

PLUNGE

m m

in.

OVERHEAD WELDING CURRE TIME LIFT NT, A s cycl m in. es m

1. 6 1. 6 1. 6 1. 6 1. 6 1. 6 2. 4 2. 4 3. 2 3. 2

0.0 62 0.0 62 0.0 62 0.0 62 0.0 62 0.0 62 0.0 93 0.0 93 0.1 25 0.1 25

3. 2 3. 2 3. 2 3. 2 3. 2 3. 2 4. 7 4. 7 6. 4 6. 4

0.1 25 0.1 25 0.1 25 0.1 25 0.1 25 0.1 25 0.1 87 0.1 87 0.2 50 0.2 50

300

2. 4 2. 4 2. 4 3. 2

0.0 93 0.0 93 0.0 93 0.1 25

3. 2 3. 2 3. 2 3. 2

0.1 25 0.1 25 0.1 25 0.1 25

150

450 500 550 675 800 1200 1500 1700 2050

200 250 325

0.1 5 0.1 5 0.2 5 0.3 3 0.4 2 0.5 5 0.6 7 0.8 4 1.0 0 1.4 0

8-9

0.2 5 0.4 0 0.5 0 0.6 5

15

1012 15 20 25 33 40 5055 65 72

24 30 39

PLUNGE

m m

in.

VERTICAL WELDING CURRE TIME LIFT NT, A s cycl m in. es m

1. 6 1. 6 1. 6 1. 6 1. 6 1. 6 2. 4 2. 4 3. 2 3. 2

0.0 62 0.0 62 0.0 62 0.0 62 0.0 62 0.0 62 0.0 93 0.0 93 0.1 25 0.1 25

3. 2 3. 2 3. 2 3. 2 3. 2 3. 2 4. 7 4. 7 6. 4 6. 4

0.1 25 0.1 25 0.1 25 0.1 25 0.1 25 0.1 25 0.1 87 0.1 87 0.2 50 0.2 50

300

0.1 8-9 1. 0.0 5 6 62 450 0.1 10- 1. 0.0 7 12 6 62 500 0.2 15 1. 0.0 5 6 62 600 0.3 20 1. 0.0 3 6 62 750 0.3 20 1. 0.0 3 6 62 875 0.4 28 1. 0.0 7 6 62 1275 0.6 36 1. 0.0 0 6 62 1700 0.7 50 2. 0.0 3 4 93 NOT RECOMMENDED

2. 4 2. 4 2. 4 2. 4

0.0 93 0.0 93 0.0 93 0.0 93

3. 2 3. 2 3. 2 3. 2

0.1 25 0.1 25 0.1 25 0.1 25

180

PLUNGE

m m

in.

3. 2 3. 2 3. 2 3. 2 3. 2 3. 2 4. 7 4. 7

0.125

3. 2 3. 2 3. 2 3. 2

0.125

0.125 0.125 0.125 0.125 0.125 0.187 (A)

0.187 (A)

NOT RECOMMENDED

225 275 350

0.2 0 0.3 0 0.4 0 0.6 0

12 18 24 36

2. 4 2. 4 2. 4 3. 2

0.0 93 0.0 93 0.0 93 0.1 25

0.125 0.125 0.125

11. 1 12. 7

0.4 37 0.5 00

3. 2 3. 2

0.12 400 5 0.12 460 5

0.8 0 0.9 0

48 54

3. 2 3. 2

0.1 25 0.1 25

3. 9 4. 7

0.1 56 0.1 87

400 460

0.8 0 0.9 0

48 54

3. 9 4. 7

0.1 56 0.1 87

3. 9 4. 7

0.1 56 0.1 87

(A) STUD ARC WELDED ALUMINUM STUDS REQUIRE INERT GAS SHIELDING: SEE TABLE 2.

430 475

0.7 0 0.8 0

42 48

3. 2 3. 2

0.1 25 0.1 25

3. 9 4. 7

0.156 0.187

There is a range of setting combinations for all stud diameters. The same total energy input can be obtained by varying current input and arc time. For example, a low current input can be compensated for, to some extent, by increasing weld time. Table 3 also shows typical weld settings for various SW stud diameters. The ideal settings in each weld position for a particular stud and workpiece should be established by beginning with typical settings and then varying them within the allowable range to meet the required conditions. Typical settings are based on good weld conditions, such as a clean workpiece, good ground, and others. Also necessary is a proper dc power source, which should have the following characteristics: • • • •

HIGH OPEN-CIRCUIT VOLTAGE (70 TO 100 V) RAPID OUTPUT CURRENT RISE DROOPING OUTPUT VOLT-AMPERE CURVE HIGH CURRENT OUTPUT OVER A SHORT TIME

Stud arc welding requires a very short weld time (Table 3), but a very high current input, when compared with other types of arc welding. Usually, the duty cycle for this process is much lower than that of other welding processes. Constantvoltage types of dc power sources are not suitable for stud welding. Many of the available dc power sources are acceptable for stud welding, including those supplied as integrated power/control systems by the stud-welding manufacturer. These integrated systems use both three-phase and single-phase incoming alternating current (ac) power. Both standard and special voltage units are available. Single-phase units are relatively low cost, portable systems that are usually suitable for smaller-diameter ( ≤ 12.7 mm, or

1 in.) stud arc welding. Three-phase units are preferred for larger2

diameter studs, because they provide a balanced load on incoming power lines and have smoother arc characteristics. A transformer/rectifier power source, combined with an integral stud-welding control system, is shown in Fig. 2. Weld cable length and cable size also can affect the total output of a power source. Basically, a larger welding cable size minimizes current loss that is due to cable length, resistance, and heating. Thus, when the distance from the power source to the gun increases significantly, or when the number and size of studs applied per minute are large, larger welding cables should be used. Smaller dc power sources suitable for stud arc welding can be wired in parallel to produce the necessary current range. Ferrules, or ceramic arc shields, are used in most stud arc weld applications. They are available in a wide variety of sizes

and shapes to fit specific stud base designs and applications. For example, there are specific designs for welding studs through metal deck, to the fillet or heel of an angle, to round pipes or bars, or in a vertical position. A ferrule is placed over every stud at the weld end, where it is held in place by a grip on the stud-welding gun. The ferrule is used only once and must be removed from the stud after welding is completed to allow the inspection of the completed weld. Typical ferrule configurations are shown in Fig. 8.

FIG. 8 CROSS SECTIONS OF FERRULES SHOWING SOME OF THE MANY VARIETIES AVAILABLE FOR DIFFERENT STUD BASE GEOMETRIES, WELD POSITIONS, AND APPLICATIONS

Ferrule design is important, because it controls certain functions during welding. For example: •

• • •

VENTS ON THE BOTTOM OF THE FERRULE ALLOW WELD GASES TO ESCAPE FROM THE WELD AREA AND RESTRICT THE INFLOW OF AIR TO MINIMIZE WELD POROSITY AND OXIDATION. THE INTERNAL CAVITY OF THE FERRULE CONFINES AND SHAPES THE MOLTEN METAL INTO A FLASH (FILLET) AROUND THE STUD PERIPHERY. HEAT IN THE WELD AREA IS CONCENTRATED AND CONTAINED. FLASH AND WELD SPATTER ARE MINIMIZED.

Although the dimensions of the flash are controlled by the ferrule configuration, the flash diameter and height must be taken into consideration when designing mating parts. Manufacturer specifications on finished stud flash dimensions should be followed, and test welds should be made to establish part fit. Flash dimensions can be accommodated by any of the five methods shown in Fig. 9.

FIG. 9 COMMON METHODS FOR ACCOMMODATING WELD FLASH (FILLET) WITH STUD ARC WELDS. TYPICAL DIMENSIONS FOR A, B, AND C ARE GIVEN IN TABLE 4.

The typical dimensions for flash clearance are shown in Table 4. It should be noted that the flash (fillet) of a stud weld is not the same as the fillet that is produced by conventional welding techniques. It may have areas of nonfusion on its vertical leg or shrink fissures, which have no adverse affect on weld strength or ductility (Ref 4). Consequently, it is not subject to the profile and inspection criteria used for a conventional weld fillet.

TABLE 4 DIMENSIONS FOR COUNTERBORE AND COUNTERSINK WELD FLASH CLEARANCE A, B, and C are shown in Fig. 9.

STUD BASE DIAMETER COUNTERBORE A B mm in. mm in. mm in. 1 6.4 11.1 0.437 3.2 0.125 7.9 9.5 11.1 12.7 15.9

4 5 16 3 8 7 16 1 2 5 8

90° COUNTERSINK C mm in. 3.2 0.125

12.7 0.500 3.2

0.125 3.2

0.125

15.1 0.593 3.2

0.125 3.2

0.125

16.7 0.656 4.7

0.187 3.2

0.125

19.1 0.750 4.7

0.187 4.7

0.187

22.2 0.875 5.5

0.218 4.7

0.187

19.1

3 4

28.6 1.125 7.9

0.312 4.7

0.187

Note: Dimensions can vary, depending on stud style and ceramic arc shield selected. Consult manufacturer for details. Base Plate Material and Thickness. To obtain full strength, the stud arc welding of low-carbon steel, stainless steel, or aluminum studs should be made to a base material of sufficient thickness so that the stud fails in tension or torque, rather than by pulling a hole in the base-metal workpiece. In the case of stainless or low-carbon steel studs, the base-metal thickness should be at least one-third that of the stud diameter. For aluminum studs, the aluminum base plate material should be at least half that of the stud diameter. Thinner base plates can be used if strength is not the primary design characteristic that is required. On thinner materials, the studs will pull a hole in the plate at strength levels below their maximum. If the plate is too thin, then the stud will bum a hole through it. The recommended base plate thickness for fullstrength development and the minimum thicknesses required to prevent burning through the base plate are given in Table 5.

TABLE 5 RECOMMENDED BASE METAL THICKNESSES FOR STUD ARC WELDING STUD BASE DIAMETER

STEEL BASE METAL FOR FULL STRENGTH(A)

MINIMUM(B)

mm 4.8

mm 1.59

in. 0.062

mm in. mm 0.91 0.036 3.18

in. 0.125

mm 3.18

in. 0.125

3.18

0.125

1.21 0.048 3.18

0.125

3.18

0.125

3.18

0.125

1.52 0.060 4.76

0.187

3.18

0.125

4.76

0.187

1.90 0.075 4.76

0.187

4.76

0.187

4.76

0.187

2.28 0.090 6.35

0.250

4.76

0.187

4.76

0.187

3.04 0.120 6.35

0.250

6.35

0.250

6.35

0.250

3.68 0.145 . . .

...

...

...

7.94

0.318

4.70 0.185 . . .

...

...

...

9.53

0.375

6.35 0.250 . . .

...

...

...

9.53

0.375

9.53 0.375 . . .

...

...

...

6.4 7.9 9.5 11.1 12.7 15.9 19.1 22.2 25.4

in. 3 16 1 4 5 16 3 8 7 16 1 2 5 8 3 4 7 8

1

ALUMINUM BASE METAL WITHOUT BACKUP MINIMUM WITH BACKUP(C)

(A) NEAREST COMMONLY AVAILABLE BASE METAL THICKNESS. (B) STEEL WITHOUT BACKUP--WILL NOT DEVELOP FULL STUD STRENGTH. (C) METAL BACKUP REQUIRED TO PREVENT MELT-THROUGH OF BASE METAL The most widely used combinations of base plate and stud materials are shown in Table 6. With low-carbon steel studs, no preheat or postheat treatments are needed when welding to low-carbon steel or austenitic stainless steel base plate materials. As carbon content in the base plate increases into the medium range, heat treatment for stud welding may be required. Welding to high-carbon base plate is not suggested. When base plate/stud combinations are questionable, an application qualification test is suggested (Ref 5). Details on this procedure are provided in the section "Stud-Welding Quality Control, Qualification, and Inspection" in this article. For best results when welding stainless steel studs to lowcarbon steel plate (especially when working stress levels are very high or there are repetitive loading cycles), the studs should either be manufactured from annealed-in-process materials or postmanufacture annealed before use.

TABLE 6 TYPICAL COMBINATIONS OF BASE AND STUD METALS FOR STUD ARC WELDING

BASE METAL LOW-CARBON STEEL, AISI 1006 TO 1022(A) STAINLESS STEEL, 300 SERIES(B), 405, 410, AND 430 ALUMINUM ALLOYS, 5000 SERIES(C)

STUD METAL LOW-CARBON STEEL, AISI 1006 TO 1020; STAINLESS STEEL, 300 SERIES LOW-CARBON STEEL, AISI 1006 TO 1020; STAINLESS STEEL, 300 SERIES ALUMINUM ALLOYS, 5000 SERIES(C)

Source: Ref 1

(A) REFER TO ANSI/AWS D1.1-92, TABLE 4.1 (GROUPS I AND II) FOR APPROVED STEELS. (B) EXCEPT FOR THE FREE-MACHINING TYPE 303 STAINLESS STEEL. (C) REFER TO ANSI/AWS D1.2-89, TABLES 7.1 AND 7.4. Good welding practice requires that the base plate material at the weld spot be cleaned. Common contaminants that can result in unsatisfactory welds if not removed include paint, heavy mill scale, heavy rust, oxidation, oil or grease, and plating, such as galvanizing and anodizing on aluminum. Removal methods vary according to the contaminant, the base material, and the end use. Wire brushing and grinding may be satisfactory on heavy structural steel but can be destructive to thin-gage metals or anodized aluminum, because they may reduce the thickness to a point where bum through or lessthan-full-strength welds result. Suggested cleaning methods include grinding, wire brushing, or needle scaling on heavy materials; using solvents to clean grease and oil; and using a noncontaminating stainless wire brush or a nonmetallic foam disk to clean aluminum and milling. Because anodizing is difficult to remove from aluminum without reducing the metal thickness, consideration should be given to anodizing the base plate and stud after the welding operation. Similarly, the weld end of the stud should be free from materials that would contaminate the weld, including paint, rust, galvanizing, and others. Studs cannot be coated with any nonconductive material that would interfere with the flow of welding current. Light copper flashing and nickel or chrome plating usually do not cause welding problems. Because a high-amperage, short-time current is typical in the stud arc welding process, it is also important that the grounding spot(s) on the base material be clean, that the ground be tight, and that all cables be in good condition and have tight connections. The position of the ground(s) also can influence stud-welding quality. This depends on the geometry of the base-metal shape and is the result of electromagnetic arc deflection, or arc blow (Ref 6). Occurrences are more common when studs are very near a free edge or are on long and narrow base plates, hollow pipe sections, or irregular peripheries, and can be corrected by moving the ground to a central position or using multiple grounds. Because current flow is usually away from the ground toward the heavier or larger area of the base plate, arc blow is characterized by a lack of flash or fillet on the side of the stud nearer the edge, end, or smaller area. Stud Strength. Fasteners will develop full material strength when they have been stud arc welded to compatible base plate alloys. Tables 7, 8, and 9 show the tension and torque loads for various threaded stud diameters, based on minimum specified stud strengths (Ref 2, 3, and 6). For unthreaded stud fasteners, the yield strength and tensile or ultimate strength is calculated by:

YIELD STRENGTH = ASFY TENSILE STRENGTH = ASFU where As is the area of the stud shank (in.2), Fy is the specified stud material yield strength (minimum psi) and Fu is the specified stud material ultimate strength (minimum psi).

TABLE 7 MECHANICAL PROPERTIES OF LOW-CARBON STEEL STUD ARC WELDED FASTENERS Fasteners have 380 MPa (55 ksi) minimum ultimate strength and 345 MPa (50 ksi) minimum yield strength. STUD THREAD DIAMETER(A)

MEAN EFFECTIVE THREAD AREA(B)

YIELD TENSILE LOAD(C)

ULTIMATE TENSILE LOAD

YIELD TORQUE(D)

mm2 in.2

kn

lbf

kn

lbf

j

ULTIMATE TORQUE

ULTIMATE SHEAR LOAD(E)

10-24 UNC 10-32 UNF

11

0.017 3.8

850

4.2

935

3.6

ft · j lbf 32(F) 4.0

ft · lbf kn

lbf

35.1(F) 3.1

701

13

0.020 4.4

1,000

4.9

1,100

4.3

38(F)

4.7

41.3(F) 3.7

825

1 -20 4

21

0.032 7.1

1,600

7.8

1,760

9.1

6.7

9.9

7.3

5.9

1,320

23

0.036 8.0

1,800

8.8

1,980

10.2

7.5

11.3

8.3

6.6

1,485

34

0.052 11.6

2,600

12.7

2,860

18.4

13.6

20.2

14.9

9.5

2,145

37

0.058 12.9

2,900

14.2

3,190

20.5

15.1

22.5

16.6

10.6

2,393

50

0.078 17.3

3,900

19.1

4,290

33.1

24.4

36.3

26.8

14.3

3,218

57

0.088 19.6

4,400

21.5

4,840

37.3

27.5

41.1

30.3

16.1

3,630

68

0.106 23.6

5,300

25.9

5,830

52.3

38.6

57.6

42.5

19.5

4,373

76

0.118 26.2

5,900

28.9

6,490

58.3

43.0

64.1

47.3

21.7

4,868

92

0.142 31.6

7,100

34.7

7,810

80.3

59.2

88.3

65.1

26.0

5,856

103

0.160 35.6

8,000

39.1

8,800

90.4

66.7

99.4

73.3

29.4

6,600

146

0.226 50.3

11,300 55.3

41.5

9,323

UNC 1 -28 4

UNF 5 -18 16

UNC 5 -24 16

UNF 3 -16 8

UNC 3 -24 8

UNF 7 -14 16

UNC 7 -20 16

UNF 1 -13 2

UNC 1 -20 2

UNF 5 -11 8

12,430 159.6 117.7 175.5 129.5

UNC 5 -18 8

165

0.255 56.7

12,750 62.4

14,025 180.1 132.8 198.1 146.1

46.8

10,519

215

0.334 74.3

16,700 81.7

18,370 283.1 208.8 311.3 229.6

61.3

13,778

240

0.372 82.7

18,600 91.0

20,460 315.2 232.5 346.8 255.8

68.3

15,345

298

0.462 102.8 23,100 112.8 25,355 456.8 336.9 501.4 369.8

84.6

19,017

328

0.509 113.2 25,450 124.5 27,995 503.1 371.1 553.6 408.3

93.4

20,996

91 437

0.606 134.8 30,300 148.0 33,275 684.7 505.0 751.9 554.6 0.678 150.8 33,900 165.9 37,290 766.0 565.0 842.6 621.5

111.0 24,956 124.4 27,967

UNF 3 -10 4

UNC 3 -16 4

UNF 7 -9 UNC 8 7 -14 8

UNF 1-8 UNC 1-14 UNF

(A) UNC, UNIFIED COARSE THREAD SERIES; UNF, UNIFIED FINE THREAD SERIES. (B) MEAN EFFECTIVE THREAD AREA IS BASED ON A MEAN FULL DIAMETER MIDWAY BETWEEN MINOR AND PITCH THREAD DIAMETERS. (C) IN PRACTICE, A STUD SHOULD NOT BE USED AT OR HIGHER THAN YIELD LOAD. A FACTOR OF SAFETY SHOULD BE APPLIED, AND 60% OF YIELD IS COMMONLY USED, ALTHOUGH OTHER VALUES MAY BE USED AT DISCRETION OF USER. (D) TORQUE FIGURES ARE BASED ON THE ASSUMPTION THAT EXCESSIVE THREAD DEFORMATION HAS NOT AFFECTED THE PROPORTIONAL RANGE OF TORQUE/TENSION RELATIONSHIP. AN AVERAGE TORQUE COEFFICIENT OF 0.20 WAS USED IN THESE CALCULATIONS. (E) SHEAR LOAD IS BASED UPON 0.75 TIMES ULTIMATE LOAD. THE USER SHOULD APPLY AN APPROPRIATE SAFETY FACTOR TO THESE FIGURES. (F) VALUE GIVEN IN IN. · LBF. TABLE 8 MECHANICAL PROPERTIES OF STAINLESS STEEL STUD ARC WELDED FASTENERS Fasteners have 520 MPa (75 ksi) minimum ultimate strength and 205 MPa (30 ksi) minimum yield strength. STUD THREAD DIAMETER(A)

MEAN EFFECTIVE THREAD AREA(B)

YIELD TENSILE LOAD(C)

ULTIMATE TENSILE LOAD

YIELD TORQUE(D)

mm2

kn

kn

j

in.2

lbf

lbf

ULTIMATE TORQUE

ft · lbf (F)

j

5.4

ULTIMATE SHEAR LOAD(E)

ft · lbf (F)

47.8

kn

lbf

4.3

956

10-24 UNC 10-32 UNF

11

0.017 2.3

510

5.7

1,275

2.2

19.1

13

0.020 2.7

600

6.7

1,500

2.5

22.5(F) 6.4

56.3(F) 5.0

1,125

1 -20 4

21

0.032 4.3

960

10.7

2,400

5.4

4.0

13.6

10.0

8.0

1,800

23

0.036 4.8

1,080

12.0

2,700

6.1

4.5

15.3

11.3

9.0

2,025

34

0.052 6.9

1,560

17.3

3,900

11.0

8.1

27.5

20.3

13.0

2,925

UNC 1 -28 4

UNF 5 -18 16

UNC

5 -24 16

37

0.058 7.7

1,740

19.3

4,350

12.3

9.1

30.6

22.6

14.5

3,263

50

0.078 10.4 2,340

26.0

5,850

19.8

14.6

49.6

36.6

19.5

4,388

57

0.088 11.7 2,640

29.4

6,600

22.4

16.5

56.0

41.3

22.0

4,950

68

0.106 14.1 3,180

35.4

7,950

31.5

23.2

78.6

58.0

26.5

5,963

76

0.118 15.7 3,540

39.4

8,850

35.0

25.8

87.4

64.5

29.5

6,638

92

0.142 18.9 4,260

47.4

10,650 48.1

35.5

120.4

88.8

35.5

7,988

103

0.160 21.4 4,800

53.4

12,000 54.2

40.0

135.6

100.0

40.0

9,000

146

0.226 30.2 6,780

75.4

16,950 95.7

70.6

239.4

176.6

56.5

12,713

165

0.255 34.0 7,650

85.1

19,125 108.1 79.7

270.1

199.2

63.8

14,344

215

0.334 44.6 10,020 111.4 25,050 169.9 125.3

424.50 313.1

83.6

18,788

240

0.372 49.6 11,160 124.1 27,900 189.1 139.5

472.9

348.8

93.1

20,925

298

0.462 61.7 13,860 154.1 34,650 274.0 202.1

685.1

505.3

115.6 25,988

328

0.509 67.9 15,270 169.8 38,175 301.9 222.7

754.8

556.7

127.4 28,631

391 437

0.606 80.9 18,180 202.2 45,450 410.8 303.0 0.678 90.5 20,340 226.2 50,850 459.6 339.0

1027.0 757.5 1149.0 847.5

151.6 34,088 169.6 38,138

UNF 3 -16 8

UNC 3 -24 8

UNF 7 -14 16

UNC 7 -20 16

UNF 1 -13 2

UNC 1 -20 2

UNF 5 -11 8

UNC 5 -18 8

UNF 3 -10 4

UNC 3 -16 4

UNF 7 -9 UNC 8 7 -14 8

UNF 1-8 UNC 1-14 UNF

(A) UNC, UNIFIED COARSE THREAD SERIES; UNF, UNIFIED FINE THREAD SERIES. (B) META, MEAN EFFECTIVE THREAD AREA, IS BASED ON A MEAN FULL DIAMETER MIDWAY BETWEEN MINOR AND PITCH THREAD DIAMETERS. (C) IN PRACTICE, A STUD SHOULD NOT BE USED AT OR HIGHER THAN YIELD LOAD. A FACTOR OF SAFETY SHOULD BE APPLIED, AND 60% OF YIELD IS COMMONLY USED, ALTHOUGH OTHER VALUES MAY BE USED AT THE DISCRETION OF USER. (D) TORQUE FIGURES ARE BASED ON THE ASSUMPTION THAT EXCESSIVE THREAD DEFORMATION HAS NOT AFFECTED THE PROPORTIONAL RANGE OF TORQUE/TENSION RELATIONSHIP. AN AVERAGE TORQUE COEFFICIENT OF 0.20 WAS USED IN THESE CALCULATIONS. (E) SHEAR LOAD IS BASED UPON 0.75 TIMES ULTIMATE LOAD. THE USER SHOULD APPLY AN APPROPRIATE SAFETY FACTOR TO THESE FIGURES.

(F) VALUE GIVEN IN IN. · LBF. TABLE 9 MECHANICAL PROPERTIES OF ALUMINUM STUD ARC WELDED FASTENERS Fasteners have 290 MPa (42 ksi) minimum ultimate strength and 205 MPa (30 ksi) minimum yield strength. STUD THREAD DIAMETER(A)

10-24 UNC 10-32 UNF 1 -20 UNC 4 1 -28 UNF 4 5 -18 UNC 16 5 -24 UNF 16 3 -16 UNC 8 3 -24 UNF 8 7 -14 UNC 16 7 -20 UNF 16 1 -13 UNC 2 1 -20 UNF 2

MEAN EFFECTIVE THREAD AREA(B)

YIELD TENSILE LOAD(C)

ULTIMATE TENSILE LOAD

YIELD TORQUE(D)

ULTIMATE TORQUE

ULTIMATE SHEAR LOAD(E)

mm2 11

in.2 0.017

kn 2.3

lbf 510

kn 3.2

lbf 714

j 2.2

13 21

0.020 0.032

2.7 4.3

600 960

3.9 6.0

840 1344

23

0.036

4.8

1080

6.7

34

0.052

6.9

1560

37

0.058

7.7

50

0.078

57

j 3.0

ft · lbf 26.8(F)

kn 1.9

lbf 428

2.5 5.4

ft · lbf 19. 1(F) 22.5(F) 4.0

3.6 7.6

31.5(F) 5.6

2.2 3.6

504 806

1512

6.1

4.5

8.5

6.3

4.0

907

9.7

2184

11.0 8.1

15.5

11.4

5.8

1310

1740

10.8

2436

12.3 9.1

17.2

12.7

6.5

1462

10.4

2340

14.6

3276

19.8 14.6

27.8

20.5

8.7

1966

0.088

11.7

2640

16.4

3696

22.3 16.5

31.3

23.1

9.6

2218

68

0.106

14.1

3180

19.8

4452

31.5 23.2

44.1

32.5

11.9

2671

76

0.118

15.7

3540

22.0

4956

35.0 25.8

48.9

36.1

13.2

2974

92

0.142

18.9

4260

26.5

5964

48.1 35.5

67.4

49.7

15.9

3578

103

0.160

21.4

4800

29.9

6720

54.2 40.0

75.9

56.0

17.9

4032

(A) UNC, UNIFIED COARSE THREAD SERIES; UNF, UNIFIED FINE THREAD SERIES. (B) MEAN EFFECTIVE THREAD AREA IS BASED ON A MEAN FULL DIAMETER MIDWAY BETWEEN MINOR AND PITCH THREAD DIAMETERS. (C) IN PRACTICE, A STUD SHOULD NOT BE USED AT OR HIGHER THAN YIELD LOAD. A FACTOR OF SAFETY SHOULD BE APPLIED, AND 60% OF YIELD IS COMMONLY USED, ALTHOUGH OTHER VALUES MAY BE USED AT DISCRETION OF USER. (D) TORQUE FIGURES ARE BASED ON THE ASSUMPTION THAT EXCESSIVE THREAD DEFORMATION HAS NOT AFFECTED THE PROPORTIONAL RANGE OF TORQUE/TENSION RELATIONSHIP. AN AVERAGE TORQUE COEFFICIENT OF 0.20 WAS USED IN THESE CALCULATIONS. (E) SHEAR LOAD IS BASED ON 0.60 TIMES ULTIMATE LOAD. THE USER SHOULD APPLY AN APPROPRIATE SAFETY FACTOR TO THESE FIGURES. (F) VALUE GIVEN IN IN. · LBF. For threaded studs, the area is based on the mean effective thread area, which is calculated by:

AS = 0.7854 [D- (0.9743/N)]2

where D is the nominal diameter of the stud and N is the number of threads per inch. Process Variations and Special Equipment. There are several specific applications that lend themselves to special

variations of the stud arc welding technique. One application is the welding of studs to thin base materials that are less than the minimum thicknesses listed in Table 5. Although full weld base strength is usually not achieved, the resulting strength level is suitable for the application loadings involved. This process variation, which is called short-cycle stud arc welding, does not use a ceramic ferrule. Instead, higher weld current is used with very short times, which minimizes penetration of the stud into the base plate. Normally, this variation is used with stud diameters ≤ 9.5 mm ( ≤ 0.375 in.) in situations where backside marking is not considered detrimental. The second special process is gas-shielded arc welding, which also does not use a ceramic ferrule. Instead, the studwelding area is shielded by an inert gas, usually argon. This process can be used with both steel and aluminum, but is more widely used with the latter material. The welding variables fall into a very narrow range, and conditions for application usually include tight tolerances and a tightly controlled setup. Consequently, the stud arc welding equipment most often consists of a fixed-gun production unit. A typical application is the welding of aluminum studs with special end configurations to aluminum kitchen utensils. Special equipment is frequently used with the short-cycle and/or gas-shielded arc welding processes, because the application parameters must be tightly controlled and involve large quantities of studs applied in a production environment. Equipment manufacturers assemble many types of special equipment for production-line use. The equipment can be used with automatic stud feed, automatic stud and ferrule feed, computer-programmed indexing, robotic stud-welding guns, and other mechanisms. A typical automated stud arc welding production unit is shown in Fig. 10. A unit can involve simple, column-mounted, single-gun systems or sophisticated, multigun, multiple feed units that cost thousands of dollars.

FIG. 10 TYPICAL AUTOMATIC STUD-FEED HAND-GUN SYSTEM FOR PORTABLE OPERATION USING TEMPLATES OR FIXTURES ON THE WORKPIECE

References cited in this section

1. "RECOMMENDED PRACTICES FOR STUD WELDING," ANSI/AWS C5.4, AWS 2. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," AWS 3. "STRUCTURAL WELDING CODE--ALUMINUM," ANSI/AWS D1.2, AWS 4. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," FOOTNOTE 27, AWS 5. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," PARAGRAPH 7.6, AWS 6. W.A. BAESLACK, G. FAYER, S. REAM, AND C.E. JACKSON, QUALITY CONTROL IN ARC STUD WELDING, WELD. J., NOV 1975, P 789-798 Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Fixturing and Tooling for Stud Arc Welding Regardless of the stud-welding method employed, certain "expendable" accessories are required. These include such items as a chuck or collet to hold the particular stud being welded, a ferrule grip for the fastener, a foot that holds the ferrule grip, legs that attach to the gun and adjust to accommodate various stud lengths, and other minor items. The operating life of these items is variable, depending on the care and maintenance they receive. For example, a chuck may last for 5000 to 25,000 welds. Legs last indefinitely, whereas ferrule grips and feet usually have a shorter usage life. Accessories are relatively inexpensive items that should be planned for when preparing budgetary estimates on any given project. The extent and sophistication of tooling for stud welding reflects the required production rate and the total number of studs to be welded. Locating the stud arc welding centers can be as simple as laying out the workpiece and center punching the locations either directly or through a template, as shown in Fig. 11. The studs are then placed in the punch marks, the stud-welding gun is held vertically, and the weld is initiated. Although operator skill is a factor, careful welders can achieve a perpendicularity of ±5° and a location tolerance of approximately 1.2 mm (0.046 in.). Often, the cover plate in a base plate/cover plate assembly can be used as the template for marking.

FIG. 11 CENTER PUNCHING BASE MATERIAL THROUGH A TEMPLATE

A permanent template becomes more practical when an increasing number of studs are to be welded on repetitive locations. Stud arc welded fasteners can be welded directly through temperature-resistant material which should be spaced off the work by 1.65 to 6.35 mm (0.065 to 0.250 in.) to allow expelled weld gases and weld spatter to escape without restriction, which could adversely affect weld quality. This type of template arrangement is shown in Fig. 12. Note that the template holes are drilled slightly larger than the outside dimension of the ceramic ferrule. Before the

template is prepared, it is good practice to consult the manufacturer specifications for the ceramic ferrule dimensions to be used with the stud being welded. Location accuracy, in this case, can be as tight as ±0.78 mm (±0.031 in.).

FIG. 12 WELDING THROUGH A TEMPLATE USING A FERRULE WITH STUD ARC WELDING FASTENERS

Even greater accuracy can be obtained, both in the vertical and horizontal directions, by adapting the template hole to fit a drill jig bushing tightly and by inserting the template adapter (or a ferrule tube adapter) through the bushing to the workpiece. The ferrule tube adapter also can be used for welding down into deeply drilled holes or into tight-fit areas (Fig. 13).

FIG. 13 PORTABLE STUD ARC WELDING GUN EQUIPPED WITH ACCESSORIES FOR WELDING STUDS THROUGH

OR INTO A DRILLED OR FORMED HOLE

With closely spaced studs, the foot grip is sometimes modified so that after welding the initial stud at a center-punched location, the foot can be held against the reference stud to weld the next or adjacent stud on the required spacing. The verticality of the weld stud can be ensured by several methods. The most accurate is obviously a fixed gun on a slide mechanism of the type used for automatic-feed, high-production work. This type is also the most expensive. A portable gun can be mounted on a machine slide or drill press by fabricating a bracket (Fig. 14) in cases where production quantities do not justify the expense of automatic equipment.

FIG. 14 INEXPENSIVE METHOD OF MOUNTING PORTABLE HAND GUN TO MACHINE SLIDE OR DRILL PRESS TO OBTAIN FIXED-POSITION ACCURACY

Other methods of ensuring verticality are to use a bushing-type template, a bubble level mounted on the rear of the gun, or a bipod foot arrangement, as shown in Fig. 15. With the stud and ferrule acting as one fixed point, the two bipod screws are adjusted to obtain perpendicularity. Vertical variations that are less than ±1° can be achieved with proper accessories and adjustments.

FIG. 15 USE OF AN ADJUSTABLE BIPOD FOOT ASSEMBLY TO PROVIDE THREE-POINT CONTACT, ENSURING PERPENDICULARITY OF STUD TO WORKPIECE

A wide range of accessories is available for welding studs of different lengths or for welding studs into areas with limited access. Two widely used special accessories are shown in Fig. 16 and 17. For unusual situations, the manufacturer should be consulted to suggest or, possibly, design stud-welding accessories.

FIG. 16 TYPICAL OFFSET ASSEMBLY USED WHEN STUD ARC WELDED FASTENERS MUST BE WELDED BELOW AN OBSTRUCTION

FIG. 17 TYPICAL LEG-EXTENSION ACCESSORY FOR WELDING LONG STUDS (UP TO 1.83 M, OR 6 FT). NOTE THE USE OF A SPLIT FOOT AND FERRULE GRIP; THE GUN DOES NOT HAVE TO BE STRIPPED THE FULL LENGTH OF THE STUD WHEN THE STUD IS WELDED.

Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Stud-Welding Quality Control, Qualification, and Inspection Because of its broad use for many years and the relative simplicity of its applications, stud welding in the downhand position is considered to be a prequalified procedure, that is, it is only necessary to weld two studs satisfactorily in the downhand position to qualify both the process and the operator. The two qualification studs can be bend tested, torque tested, or tension tested. The physical test is satisfactory if no failure occurs in the weld or the heat-affected zone. Stud-welding code requirements (for example, Ref 2 and 7) define the necessary tests for prequalified stud welding to the flat or down-hand position, as well as other positions. The limit on the flat position is defined as 0 to 15° slope on the

surface to which the stud is applied. Beyond 15°, a ten-stud test is required. Again, each of the ten studs in the nonflat position must be bend, tensile, or torque tested. Bend testing is 90° from the original axis, whereas torque and tension testing are conducted until failure. Failure must occur in the stud shank or in the workpiece material, not in the weld or the heat-affected zone. All failures require retesting. Satisfactory testing qualifies the welder and the process. Typical bend, torque, and tensile testing setups are shown in Fig. 18, 19, and 20. Qualification records, including stud drawings, ceramic arc shield drawings (if used), weld settings, and position should be kept.

FIG. 18 TYPICAL BEND TESTING SETUP. STUDS THAT MUST UNDERGO APPLICATION QUALIFICATION SHOULD BE BENT 90° WITHOUT WELD FAILURE (15° FOR ALUMINUM STUDS). TESTING FOR PREPRODUCTION OR TESTING DURING PRODUCTION FOR INSPECTION PURPOSES REQUIRES BEND TESTING TO 30° FROM ORIGINAL POSITION (15° FOR ALUMINUM STUDS). FOR THREADED STUDS, TORQUE TESTING PER FIG. 19 SHOULD BE USED. SOURCE: REF 2

FIG. 19 TYPICAL TORQUE TESTING ARRANGEMENT. STUDS THAT MUST UNDERGO APPLICATION QUALIFICATION OR PREPRODUCTION TESTING SHOULD BE TORQUED TO DESTRUCTION WITHOUT FAILURE IN THE WELD. PROOF TORQUE TESTING DURING INSPECTION OF THREADED STUDS SHOULD BE APPLIED AT 60 TO 66% OF THE YIELD TORQUE LOAD FOR THE STUD SIZE TESTING. REFER TO TABLES 7, 8, AND 9. SOURCE: REF 2

FIG. 20 TYPICAL TENSION TEST FIXTURE. FOR QUALIFICATION OF APPLICATION OR PREPRODUCTION TESTING, EITHER BEND OR TORQUE TESTING CAN BE SUBSTITUTED FOR TENSION TESTING.

Either two-stud or ten-stud testing should be conducted on a production piece or on a separate piece of the same type and thickness (±25%) of material that will be used in the application, in the same position in which the studs will actually be applied. In addition to physical testing, the studs that are welded for qualification also must be visually inspected, whether there are two or ten. Visual inspection consists of verifying that a full 360 ° weld flash (fillet) is formed around the stud base. There should be no gaps or undercuts in the flash or in the wetted area. As mentioned previously, flash can be irregular in both height and evenness and it can contain shrinkage fissures or cracks. Because flash is expelled, not deposited, metal, it should not be subject to the usual inspection criteria for a weld fillet. The final visual inspection should be a measurement of the after-weld length. With stud arc welding, the after-weld length of the stud will be shorter by the lengths shown in Table 3. A consistent after-weld length, combined with acceptable visual inspection and physical testing, are assurances of satisfactory welds. Figure 21 shows typical stud arc weld appearances for both acceptable and problematic welds.

FIG. 21 STUD ARC WELD CHARACTERISTICS FOR SATISFACTORY AND UNSATISFACTORY WELDS. (A) SATISFACTORY; GOOD FLASH FORMATION, WETTING ENTIRE PERIPHERY OF STUD BASE. (B) UNSATISFACTORY; PLUNGE TOO SHORT, NOT ENOUGH MATERIAL ALLOWED FOR BURN-OFF. STUD MUST PROJECT AN ADEQUATE LENGTH BEYOND THE FERRULE. THIS CONDITION CAN ALSO BE CAUSED BY BAD GROUNDING PRACTICE. (C) UNSATISFACTORY; HANG UP CAUSED BY INADEQUATE ARC LENGTH LIFT OR BY MISALIGNMENT OF STUD, RESTRICTING FREE MOVEMENT. (D) UNSATISFACTORY; MISALIGNMENT OF STUD BY POOR GUN POSITION. (E) UNSATISFACTORY; COLD WELD, RESULTING IN NONUNIFORM OR INCOMPLETE FLASH AROUND STUD PERIPHERY. (F) UNSATISFACTORY; HOT WELD, WHERE MOLTEN MATERIAL IS EXPELLED VIOLENTLY FROM WELD, RESULTING IN POOR FLASH FORMATION.

Passing the two-stud preproduction or the ten-stud application qualification test approves the procedure and operator. Production stud welding is then allowed to proceed. However, at the start of each new production period, a two-stud test is again required before production can begin. Similarly, a change in stud diameter, equipment, settings, or operator requires requalification.

During production, welding inspection and testing should be used on a continuous basis to confirm acceptable quality. This can be done on a reasonable number of studs, depending on the welding conditions and stud appearance. After the visual inspection of all stud welds, physical testing of at least one weld per hundred is usually suggested. Any weld that does not show a full 360° flash or wetness should be tested. Welds without a full flash can be repair welded, but repaired welds are also required to be physically tested. In the case of unthreaded studs, the quality inspection test can be to bend 15°. The studs can be straightened after testing, if required. Threaded studs can be subjected to a torque test to approximately 80% of the yield torque load, rather than a bend test. Failure in the weld or heat-affected zone is cause to reject the weld being tested. Failure also requires that the quantity of studs welded from the previous test to the present test be carefully inspected for weld deficiencies. Additional tests may be necessary to confirm weld quality. If the quality inspection test studs must be straightened for end use, then they should be returned to the vertical position by using a bending tool. Continuous slow pressure should be applied during the straightening process. Temperature at the time of stud welding and during stud testing is an influential parameter. Stud welding should not be conducted at temperatures below -18 °C (0 °F). When the temperature is below 0 °C (32 °F), one additional stud per hundred welded should be tested to verify weld quality. Below 10 °C (50 °F), welded studs should be tested using the bending tool, rather than by hammer testing. In colder welding temperature conditions, threaded studs should always be tested by torque, rather than by bending. Impact testing at low temperatures produces brittle failures. This is not a function of weld quality, but of the ambient temperature at the time of welding and testing. Data resulting from extensive tests conducted on studs welded and tested at various temperatures have shown that it is the temperature and manner of testing that influences the failure mode, rather than the temperature at the time of welding (Ref 8). Another factor, besides temperature, that can affect weld quality is cleanliness of the weld area. Although stud arc welding does provide some cleaning action during welding, it is not sufficient to remove heavy contaminants. Coatings or rust on studs also can cause poor welds. Plating, except for a copper flash, nickel plating, or chromium plating, should be removed from the stud end prior to welding. This includes such materials as paint, zinc, and cadmium. Stud arc welded fasteners are usually capped during plating so that the weld end is left clean. The plating or contaminant removal methods should be compatible with the base-material thickness and composition. For example, a heavy paint or rust cover on thick structural steel can be removed by grinding, needle scaling, sand blasting, or chipping, and the weld surface will still be acceptable for stud welding. Stainless steel, brass, aluminum, or other polished surfaces should be cleaned with appropriate solvents or other cleaning methods that will not destroy the material finish or contaminate the weld area. For example, a carbon-steel wire wheel or brush should not be used to clean stainless steel or aluminum base material. Studs and ferrules should be kept clean and dry prior to welding. Although cold conditions will not adversely affect either, they should be protected from exposure to adverse weather conditions, high humidity, marine atmospheres, and other extremes. When ferrules absorb water, the steam generated during the welding process can cause very poor welds and excessive weld spatter. Ferrules exposed to moisture should be oven dried at a temperature between 101 and 120 °C (215 and 250 °F) for 2 h before use. The condition of weld cables, connections, and ground clamps can affect weld quality. These items should be checked frequently and either repaired or replaced, so that the cables are not frayed, connections are tight and clean, and ground clamps are properly functioning. Ground clamps should be tightly connected to the workpiece or work platen and attached to areas that have been cleaned of all deleterious materials that would prevent good current flow.

References cited in this section

2. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," AWS 7. "WELDED STEEL CONSTRUCTION (METAL ARC WELDING), STANDARD W59," CANADIAN STANDARDS ASSOCIATION 8. KENNEDY AND D.J. LAURIE, STUD WELDING AT LOW TEMPERATURES, CAN. J. CIVIL ENG., VOL 7, 1980, P 442-455

Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Stud-Welding Safety Precautions Any welding process can be dangerous if proper safety precautions are not followed. Equipment should be properly installed according to the directions of the manufacturer and maintained in good condition. Operators should be thoroughly trained in the use of the stud-welding process. They also should be familiar with the installation, operation, and maintenance procedures for the equipment being used. Electrical shock is a potential cause of injury or death. Stud-welding systems should be properly installed and grounded using electrical connections made according to national (Ref 9) and local code guidelines. Neither operators nor anyone else in the welding area should ever touch live electrical parts, or weld in wet areas, or wear wet gloves or clothing. In addition, all workers should ensure that they are insulated from shock. All electric cables and connections should be kept in good condition. They should be inspected on a regular basis and frayed sections, broken insulation, or broken connectors should be repaired or replaced at once. Stud welding frequently uses long cable lengths, particularly on construction sites, where the cables are subject to abuse by towmotors, trucks, material storage, or foot traffic. It should be ensured that these conditions do not result in cable damage. Although weld spatter and arc flash are minimal with the stud-welding process, they do occur and precautions should be taken. All combustible or volatile materials should be removed from the weld area so that sparks or spatter cannot reach them. Gas cylinders should be stored and secured properly and checked to ensure that they do not contact any electrical cables in the welding circuit. Proper protective clothing should be worn, including boots, aprons, and gloves, as necessary. Eye protection is always necessary. Eye glasses with spectacle frames, side shields, and lenses with shade number three absorption filters should be worn by the operator at all times (Ref 10). Helpers or workers within 1.5 m (5 ft) of the weld area also should wear clear safety glasses with side shields. Fire-suppression equipment should be available in or adjacent to the weld area for immediate use in emergencies. Ventilation of the welding area is necessary. Fumes from welding and cleaning materials, such as solvents, as well as paints, epoxies, and galvanizing or other coatings, can be harmful. Ventilation can be either forced or natural, depending on the job conditions. Material suppliers should provide material safety data sheets on any items being used in the welding area, and proper precautions for potentially dangerous contents should be followed. Pinch points where fingers or hands can be caught should be avoided during material handling or where moving parts are involved in the welding process. The instructions of the manufacturer should be followed when maintaining and servicing the stud-welding equipment. If possible, all electric supplies should be disconnected and locked out during maintenance or troubleshooting work. Finally, certain welding operations produce elevated noise levels during the weld cycle. Operators and other workers near the equipment should use hearing protection that can adequately protect them, in accordance with Ref 11.

References cited in this section

9. "NATIONAL ELECTRICAL CODE" NFPA-70, NATIONAL FIRE PROTECTION ASSOCIATION 10. "SAFE PRACTICE FOR OCCUPATION AND EDUCATIONAL EYE AND FACE PROTECTION," Z87.1, ANSI 11. "WELDING, CUTTING, AND BRAZING," STANDARD 29CFR, PART 1910, OCCUPATIONAL SAFETY AND HEALTH ADMINISTRATION Stud Arc Welding

Harry A. Chambers, TRW Nelson Stud Welding Division

References

1. "RECOMMENDED PRACTICES FOR STUD WELDING," ANSI/AWS C5.4, AWS 2. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," AWS 3. "STRUCTURAL WELDING CODE--ALUMINUM," ANSI/AWS D1.2, AWS 4. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," FOOTNOTE 27, AWS 5. "STRUCTURAL WELDING CODE--STEEL," ANSI/AWS D1.1, SECTION 7, "STUD WELDING," PARAGRAPH 7.6, AWS 6. W.A. BAESLACK, G. FAYER, S. REAM, AND C.E. JACKSON, QUALITY CONTROL IN ARC STUD WELDING, WELD. J., NOV 1975, P 789-798 7. "WELDED STEEL CONSTRUCTION (METAL ARC WELDING), STANDARD W59," CANADIAN STANDARDS ASSOCIATION 8. KENNEDY AND D.J. LAURIE, STUD WELDING AT LOW TEMPERATURES, CAN. J. CIVIL ENG., VOL 7, 1980, P 442-455 9. "NATIONAL ELECTRICAL CODE" NFPA-70, NATIONAL FIRE PROTECTION ASSOCIATION 10. "SAFE PRACTICE FOR OCCUPATION AND EDUCATIONAL EYE AND FACE PROTECTION," Z87.1, ANSI 11. "WELDING, CUTTING, AND BRAZING," STANDARD 29CFR, PART 1910, OCCUPATIONAL SAFETY AND HEALTH ADMINISTRATION Stud Arc Welding Harry A. Chambers, TRW Nelson Stud Welding Division

Selected References

• "DESIGN AND APPLICATION HAND BOOK FOR LIGHT GAUGE METALWORKING," PUBLICATION CD-92, TRW NELSON STUD WELDING DIVISION, 1992 • "INDUSTRIAL DESIGN DATA--NELSON STUD WELDING PROCESS," PUBLICATION MB-12-86, TRW NELSON STUD WELDING DIVISION • "SAFETY IN WELDING AND CUTTING," ANSI/AWS Z49.1, AWS • T.E. SHOUP, "STUD WELDING," BULLETIN 214, WELDING RESEARCH COUNCIL, APRIL 1976 • "TOOLING TECHNIQUES FOR STUD WELDING," TRW NELSON STUD WELDING DIVISION • STUD WELDING, CHAPTER 9, WELDING HANDBOOK, 8TH ED., AWS, 1990

Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Introduction CAPACITOR DISCHARGE STUD WELDING is a stud arc welding process in which the tip of the stud melts almost instantly when energy stored in capacitors is discharged through it. The three basic modes of capacitor discharge (CD) stud welding are initial-gap welding, initial-contact welding, and drawn-arc welding. Figure 1 shows current-versus-time curves for the three process variations.

FIG. 1 TYPICAL CURRENT-VERSUS-TIME CURVES FOR THE THREE CD STUD WELDING METHODS. SOURCE: REF 1

The initial-gap mode (Fig. 2) is begun with the stud held away from the work surface by the welding head. At the beginning of the weld cycle, the stud is forced against the work surface, melting the tip, the face of the stud, and the adjoining work surface upon contact with the work surface. The weld is completed using the gun forces (i.e., spring pressure or air pressure) to plunge the stud into the molten materials, forming a strong welded bond between the stud and the work surface. The weld cycle time (Fig. 1) for this process is from 4 to 6 ms, and the penetration of the weld zone into the work surface is normally from 0.10 to 0.15 mm (0.004 to 0.006 in.).

FIG. 2 INITIAL-GAP CD STUD WELDING. SEE TEXT FOR EXPLANATION. SOURCE: REF 1

The initial-contact mode (Fig. 3) begins with the weld stud in contact with the work surface. The weld cycle is initiated

with a surge of current that disintegrates the weld tip, thus melting the stud face area and the work surface that it immediately contacts. The stud is forced into the molten material, forming a strong homogeneous weld. This process has a weld cycle time of approximately 6 ms, much like the gap process.

FIG. 3 INITIAL-CONTACT CD STUD WELDING. SEE TEXT FOR EXPLANATION. SOURCE: REF 1

The drawn-arc mode (Fig. 4) begins with the stud in contact with the work surface. When the weld cycle is initiated, a current surge is applied to the weld tip and the stud is retracted from the work surface, drawing an arc that melts the face of the stud and the work surface directly beneath it. The stud is then plunged into the molten pool of material, forming a welded connection. The weld cycle time for this process is longer than for the other two processes, and the heat-affected zone (HAZ) is thicker than it is in the preceding two processes.

FIG. 4 DRAWN-ARC CD STUD WELDING. SEE TEXT FOR EXPLANATION. SOURCE: REF 1

Advantages and Disadvantages. A major reason for using the CD stud welding process is that it provides a strong welded fastener with either a minimum or no reverse side marking of the work material. Another reason is that it is cost effective, especially for small-diameter fasteners. This process also allows fasteners to be welded to very thin sections of work material, as well as to thick sections (as thick as necessary), with reliable results. Furthermore, the process allows the welding of many dissimilar material combinations, such as aluminum studs to zinc die castings. The disadvantages of using the CD process are the limited stud diameters available and the fact that the work surface must be clean of mill scale, dirt, oxidation products, and oil.

Reference

1. R.L. O'BRIEN, ED., WELDING HANDBOOK, 8TH ED., VOL 2, AWS, 1991, P 317-323 Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Applications The CD process is normally used for fastener sizes of 4-40 through in.). Although there are limited uses of

1 1 -20, with lengths from 6.4 to 38.1 mm ( to 1.5 4 4

5 3 -18 and -16 diameter studs in special applications, only the more-common 16 8

stud diameters will be considered in this article. The work material thickness can typically range from 0.25 mm (0.010 in.) to as thick as the application requires. The determining factors for making successful welds are the stud diameter, the materials being used, the strength required, and the amount of allowable reverse side marking. See Table 1 for guidance on the expected strength parameters.

TABLE 1 STANDARD CAPACITOR DISCHARGE STUD FASTENER LOAD STRENGTHS STUD MATERIAL

STUD SIZE

LOW-CARBON COPPER-FLASHED STEEL

6-32 8-32 10-24 1 -20 4 5 -18 16 3 -16 8 6-32 8-32 10-24 1 -20 4 5 -18 16 3 -16 8 6-32 8-32 10-24 1 -20 4 5 -18 16 3 -16 8 1 -13 2 6-32 8-32 10-24 1 -20 4 5 -18 16 3 -16 8

STAINLESS STEEL (TYPE 304)

ALUMINUM ALLOY (5000 SERIES)

BRASS (70-30 AND 65-35)

MAXIMUM FASTENING TORQUE(A) n·m lbf · in. 0.7 6 1.4 12 1.6 14 4.9 43

ULTIMATE TENSILE LOAD kn lbf 2.2 500 3.4 765 4.3 960 7.8 1750

MAXIMUM SHEAR LOAD kn lbf 1.7 375 2.6 575 3.2 720 5.8 1300

8.1

72

12.9

2900

9.8

2200

11.9

106

19.1

4300

14.6

3250

1.1 2.3 2.6 8.5

10 20 23 75

3.5 5.6 6.8 12.8

790 1260 1530 2880

2.6 4.1 5.2 9.6

590 940 1150 2160

14.2

126

16.7

3750

12.5

2800

20.0

186

21.6

4850

16.0

3600

0.40 0.85 1.13 4.5

3.5 7.5 10 40

1.7 2.6 3.3 6.0

375 585 735 1360

1.0 1.6 2.0 3.8

235 365 460 850

7.9

70

10.2

2300

6.2

1400

9.1

81

15.1

3400

9.3

2100

15.8

140

25.0

5500

13.4

3000

0.9 1.8 2.09 6.9

8 16 18.5 61

2.7 3.8 4.6 8.7

600 860 1040 1950

1.7 2.5 3.0 5.7

390 560 680 1275

11.5

102

14.6

3280

9.5

2140

16.9

150

21.4

4800

14.1

3160

(A) THESE VALUES SHOULD DEVELOP FASTENER TENSION TO SLIGHTLY LESS THAN YIELD POINT AND SHOULD BE USED ONLY AS A GUIDE; BASIC SPECIFICATIONS SHOWN ABOVE COVER TYPICAL MECHANICAL PROPERTY VALUES.

The stud and workpiece materials can be common low-carbon steel, stainless steel, or aluminum. Also used are mediumcarbon steel, lead-free brass and copper, Inconel, titanium, René 41, zirconium, gold, silver, and platinum. Table 2 shows

some of the more popular combinations of materials. Various studs and plate materials can be combined, except when aluminum-base material is used.

TABLE 2 WELDING CAPABILITIES OF CAPACITOR DISCHARGE STUD FASTENERS BASE MATERIAL

MILD STEEL (1006 THROUGH 1030) MEDIUM-CARBON STEEL (1030 THROUGH 1050) GALVANIZED SHEET DUCT OR DECKING STRUCTURAL STEEL STAINLESS STEEL (TYPES 405, 410, 430, AND 300 SERIES, EXCEPT 303) LEAD-FREE BRASS, ELECTROLYTIC COPPER, LEADFREE ROLLED COPPER MOST ALUMINUM ALLOYS OF THE 1000, 3000, 5000, AND 6000 SERIES(B) DIE-CAST ZINC ALLOYS

STUD MATERIAL MILD STEEL STAINLESS (TYPES 1008, STEEL 1010) (TYPES 304, 305)

ALUMINUM (5356, 6061)

BRASS (70-30, 65-35)

EXCELLENT GOOD(A)

EXCELLENT GOOD(A)

... ...

EXCELLENT GOOD(A)

GOOD(A)

GOOD(A)

...

...

GOOD(A) EXCELLENT

GOOD(A) EXCELLENT

... ...

EXCELLENT EXCELLENT

EXCELLENT

EXCELLENT

...

EXCELLENT

...

...

EXCELLENT . . .

GOOD(A)

GOOD(A)

EXCELLENT GOOD(A)

(A) GENERALLY FULL-STRENGTH RESULTS, DEPENDING ON THE COMBINATION OF STUD SIZE AND BASE METAL. (B) OTHER MATERIALS, SUCH AS 7000 SERIES ALUMINUM, TITANIUM ALLOYS, INCONEL, AND SO ON, CAN BE WELDED UNDER SPECIFIED CONDITIONS. CD stud welding is used in many industries in a large variety of applications because it is one of the most versatile and reliable processes available. The food processing industry utilizes this process on stainless steel utensils and machinery, aluminum cookware, and appliances. The aerospace industry applies the process to aircraft engines, aircraft parts and instrument areas, and components for outer space vehicles. The appliance industry uses CD welding for its cabinets, trims, and other components. The building industry uses it for doors, facia, subassemblies, and other exposed finished areas. The insulation industry uses it in securing insulation to heating and air conditioning duct work, as well as on apparatus at power plants, schools, and other major buildings. The automotive industry uses the process to secure trim and various components to cars and trucks. The ship-building industry uses it in fabricating components in ship kitchens, in insulation applications aboard ship and in other areas where components and instruments must be secured. The electrical industry uses this process when securing components inside various controllers and equipment, as well as in trim parts for cabinets. The sign industry uses it on both structural and component members of large and small signs. Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Equipment The equipment used for CD stud welding can be either portable or stationary production-type units. The portable units consist of the basic controller that utilizes standard 110 V alternating current (ac) power input to charge the capacitors and a lightweight gun-shaped tool used for placing and welding the fasteners. The production-type units typically require

240/480 V ac three-phase incoming power and a compressed air supply. The production units are normally used for higher rates of productivity, close-tolerance work (down to ±0.13 mm, or 0.005 in.), critical reverse side marking requirements, automatic stud feeding, automatic feeding of the part to be welded, automatic location of the stud on the part, and exotic materials. The production rate of portable equipment normally ranges from 4 to 6 studs/min, depending on the application. Material handling and other factors could cause this rate to be either slower or faster. The production rates for the permanent production equipment can range from 3 to 20 studs/min, depending on the application, special requirements, and tooling. When the initial-gap or initial-contact modes of stud welding are used, welds are made on all materials without the use of shielding gases or ceramic arc shields. When using the drawn-arc mode while working with aluminum, a shielding gas would be necessary. This mode is also less flexible when working with exotic materials. Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Personnel Responsibilities The CD stud welding process is easy to operate and maintain. The operator must first ensure that his unit is connected to the proper incoming voltage. The operator then determines whether the weld gun is connected to the negative and the ground is connected to the positive. Straight polarity is used for welding in nearly all applications. The operator places a given diameter of fastener into the stud gun. The gun is then placed on the work surface. The operator closes the trigger of the gun and the welding process begins. After the welding process is completed, the operator lifts the gun off the welded stud. The operator needs to inspect the welds to ensure that the settings were correct. The controller voltage, which is set according to the recommendations of the equipment manufacturer, depends on the material and diameter of the fastener to be welded, as well as on the thickness and type of base material being used. There are also other parameters that should be considered. If it is desirable to reduce spatter and smoky conditions in the weld area, then wetting the weld area with a wetting agent just prior to making the weld is recommended. If a wetting agent is used, then it is best to add a few drops of liquid hand soap to a pint of plain water and to spray or brush the solution into the weld area. This will produce a clean weld with reduced, if not eliminated, spatter in the weld area. However, the noise of welding increases when a wet weld is made. Weld inspection is normally done by bending the fastener by applying either a torque or tensile force to it. If a tensile force is applied, then care should be given to the diameter of the hole used to secure the plate material. A hole that is as small as possible should be used to accommodate the flange of the stud. If a large hole is used, the tensile factors may be reduced, because the plate material could yield prematurely. Prior to production welding, it is recommended that sample welds be made on similar materials and tested to destruction. During production, operators should visually inspect the welds periodically. If weld inconsistency occurs, the process should cease, and the cause should be determined and eliminated. A common occurrence that contributes to inconsistency is movement of the plate material during the short weld cycle. This can occur when working with materials that are less than 3.2 mm (0.125 in.) thick. There should not be a gap between the workpiece and the backup material of the fixture or table being used. Proper fixtures and attention to their clamping system can alleviate this problem.

Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Reference

1. R.L. O'BRIEN, ED., WELDING HANDBOOK, 8TH ED., VOL 2, AWS, 1991, P 317-323 Capacitor Discharge Stud Welding Richard L. Alley, American Welding Society

Selected References

• RECOMMENDED PRACTICES FOR STUD WELDING, C5.4-84, AWS, 1984 Plasma-MIG Welding Ian D. Harris, Edison Welding Institute

Introduction PLASMA-MIG WELDING can be defined as a combination of plasma arc welding (PAW) and gas-metal arc welding (GMAW) within a single torch, where a filler wire is fed through the plasma nozzle orifice. The process can be used for both welding and surfacing. The plasma-metal inert gas (MIG) welding process was invented at the Philips Research Laboratories in Eindhoven, Netherlands, around 1969 (Ref 1). For descriptions of the PAW and GMAW processes, see the articles "Plasma Arc Welding," and "Gas-Metal Arc Welding" in this Section of the Handbook. The principles of operation, in terms of equipment, are illustrated in Fig. 1. Separate power supplies are used for the PAW and the GMAW elements of the equipment. An arc is struck between the tungsten electrode and the workpiece in a similar fashion to that of a PAW system. The filler wire can be fed to the plasma arc, either with or without the GMAW arc established. Without power supplied to the filler wire, the system can be operated as a PAW system with concentric feed of filler wire. Later versions of the system incorporated an annular electrode to replace the offset tungsten electrode in the welding torch.

FIG. 1 SCHEMATIC OF PLASMA-MIG WELDING EQUIPMENT. SOURCE: REF 1

Current and Operating Modes. The equipment can be operated either with a single power source, effectively as a PAW system with concentric filler wire feed, or with two power sources, for the plasma-MIG operation.

The polarity of the tungsten electrode is direct current, electrode negative (DCEN), as is that of the GMAW part of the system. The heat of the plasma arc is sufficient to achieve good metal transfer stability for the GMAW element, despite the fact that when this process is used separately, it is almost always used in a direct-current, electrode positive (DCEP) mode. The filler wire is heated by the constricted plasma arc, as well as by the cathode heating of its own arc, and by resistance heating along the wire extension. Therefore, the melting and deposition rates of the wire are higher than the rates achieved by heating with either arc alone. Metal transfer is governed not only by plasma streaming, but also by arc forces between the wire tip and the workpiece. Because the metal droplets are totally enclosed by the plasma stream, spray transfer takes place even though the GMAW element operates on negative polarity. Advantages and Disadvantages. The advantages of the plasma-MIG process include deposition rates and joint

completion rates that are higher than those of the conventional GMAW process. The independent control of the plasma arc and current to the filler wire leads to more control of metal deposition. This capability can yield improved productivity and good flexibility for controlling heat input and arc characteristics in both welding and surfacing operations. Good control of dilution is achieved by running the system without any power applied to the filler wire. Metal transfer stability is increased, compared to that of the conventional GMAW process, and results in lower spatter levels. The cleaning action of the plasma arc results in lower porosity in aluminum alloys, compared to that of the conventional GMAW process. Disadvantages include the capital cost of two power sources (although there are systems that are designed to operate with one), the greater complexity of the torch, and the increased maintenance time and cost associated with this complexity. With two power sources, more welding parameters need to be set up, compared to the conventional GMAW process. Equipment As noted earlier, the basic equipment includes a power source for the plasma arc and a power source for the GMAW part of the system. A special torch incorporating both a contact tip for the GMAW element and a cathode for the PAW element is required. The initial design incorporated an offset tungsten electrode, as well as a concentric conduit and

contact tip for the delivery of the consumable wire (Fig. 1). A later design incorporated a concentric cathode for the plasma arc (Fig. 2).

FIG. 2 SCHEMATIC OF MODERN PLASMA-MIG TORCH WITH ANNULAR PAW ELECTRODE AND ADDITIONAL (FOCUSING) GAS STREAM. SOURCE: REF 2

The plasma-MIG torch can be readily fitted to existing welding equipmentsuch as side beams and welding carriages, to replace the GMAW process in mechanized welding operations. Power Sources. A constant-current power source with a high-frequency circuit to initiate the pilot arc is used for the plasma arc component of the system. The power source for the GMAW component can be used as a constant-voltage or a constant-current rectifier. Power sources have welding currents that typically range from 40 to 200 A for the plasma arc and from 60 to 300 A for the GMAW element at 100% duty cycle. However, equipment with welding currents up to 800 A is available and can be used for surfacing applications. Welding Torches. A special torch with a concentric cathode for the plasma arc and a concentric conduit and contact tip

for the delivery of the consumable wire is required (Fig. 2). A water-cooled copper alloy nozzle is used to constrict the arc and to form a collimated plasma jet that exits the nozzle orifice. A plasma orifice gas and a focusing gas from the same supply are used; the latter is delivered via channels between the plasma welding electrode and the constricting nozzle. The focusing gas results in greater arc constriction and arc stability and prolongs the life of the constricting nozzle by creating a boundary gas layer between the nozzle orifice and the plasma arc. The contact tip, plasma cathode, and the constricting nozzle are all directly water cooled to provide 100% duty cycle at the typical welding currents identified in the preceding section, "Power Sources." Shielding Gases. Three shielding gases are utilized: one for the plasma (orifice) gas, one to provide additional arc

constriction and arc stability, and one for supplementary shielding. The plasma gas and the focusing gas are usually argon, because an inert gas is required to prevent oxidation of the PAW electrode. The supplementary shielding gas can be argon, argon-oxygen, argon-carbon dioxide, or argon-hydrogen, depending on the nature of the workpiece being welded or, in the case of a surfacing operation, on the material being deposited. Argon is used when welding aluminum alloys, whereas argon-oxygen and argon-carbon dioxide are used when welding steels. Argon-hydrogen is used when welding stainless steels or when surfacing with them.

References

1. W.G. ESSERS, A.C.H.G. LIEFKENS, AND G.W. TICHELAAR, PLASMA-MIG WELDING, PROC. CONF. ADVANCES IN WELDING PROCESSES, THE WELDING INSTITUTE, 1971, P 216-219 2. J.D. SWART, PLASMA-MIG BOOSTS TANK TRAILER OUTPUT, WELD. DES. FABR., FEB 1983, P 54-55, 59 Plasma-MIG Welding Ian D. Harris, Edison Welding Institute

Procedure Process Operating Procedure. The plasma arc is ignited using a pilot arc in a fashion similar to that of a PAW system. The main arc is transferred from the electrode to the workpiece and the plasma jet passes through the nozzle orifice. The system can be operated in this way, with the concentric cold wire being fed through the axis of the torch. A higher melting rate is achieved when power is applied to the wire through the contact tip and when both arcs are run simultaneously. The higher energy imparted to the wire by the plasma arc results in an increased wire deposition rate (Table 1). In this operating mode, deposition rates higher than those typical of the GMAW process can be achieved.

TABLE 1 COMPARISON OF FILLER WIRE MELTING RATES FOR PLASMA-MIG WELDING WITH AND WITHOUT THE GMAW ARC Mild steel filler wire; plasma gas, argon; shielding gas, 89% Ar + 6% CO 2 + 5% O2 PLASMA ARC CURRENT, A

110 135 160 190 190 190

PLASMA ARC VOLTAGE, V

29 30 32 34 37 38

FILLER WIRE CURRENT, A

... ... ... ... 100 150

MIG-ARC VOLTAGE, V

DIAMETER OF FILLER WIRE

MELTING RATE OF FILLER WIRE

... ... ... ... 31 32

mm 0.9 0.9 0.9 0.9 1.2 1.2

g/min 22 28 33 40 85 130

in. 0.04 0.04 0.04 0.04 0.05 0.05

oz/min 0.78 0.99 1.16 1.41 3.0 4.6

Source: Ref 1 Inspection and Weld Quality Control. Inspection requirements are similar to those of other arc welding or surfacing operations. Visual, ultrasonic, and radiographic inspection techniques are most appropriate. The dual action of the GMAW and plasma arcs results in weld quality that is sometimes higher than that achieved by the GMAW process alone. This is particularly true for aluminum alloys, because the cleaning action of the plasma arc often results in reduced porosity.

Quality control requires monitoring the welding parameters for both power sources, as well as monitoring the wire feed. In addition, the condition of the nozzle orifice (that is, the wear and concentricity of the orifice) should be monitored. Troubleshooting. The relatively complex nature of the welding torch involves increased maintenance time. Erosion of

the copper alloy nozzle orifice will cause a change in the arc shape and will affect the weld profile. Therefore, the nozzle should be checked periodically.

Reference cited in this section

1. W.G. ESSERS, A.C.H.G. LIEFKENS, AND G.W. TICHELAAR, PLASMA-MIG WELDING, PROC. CONF. ADVANCES IN WELDING PROCESSES, THE WELDING INSTITUTE, 1971, P 216-219 Plasma-MIG Welding Ian D. Harris, Edison Welding Institute

Applications Material Types. The plasma-MIG process is suitable for welding a wide variety of materials. The high heat energy

supplied by the plasma and gas-metal arcs makes the process suitable for high-melting-point materials, such as tungsten and molybdenum. The most common application is welding aluminum sheet and plate. Wear-resistant steels are used with the process in hardfacing applications. Austenitic stainless steels, such as types 308, 309, and 347, as well as nickel alloys, such as alloy 625, are used in cladding applications. Both solid and flux-cored wires (Ref 3) can be employed for welding and surfacing, although most applications involve solid wires. Industries. The plasma-MIG welding process has been used for the deposition of corrosion-resisting stainless steel and

nickel-base alloys in the offshore industry, for the general fabrication of silos and tank trailers made from aluminum alloys (Ref 2), and for hardfacing applications in the excavation equipment industry, as well as the dredging and offshore industries. Typical Components and Joints. The plasma-MIG process is suitable for welding of joints and for surfacing

operations. The wide range of heat inputs available by choosing how to apply current to the consumable electrode provides additional flexibility for surfacing operations, compared to the range of heat inputs available with an external wire feed using the GTAW/PAW or the conventional GMAW process. Single-V butt joints are commonly used for the plasma-MIG welding of plate. A full-penetration weld can be made in a single pass on a 9.5 mm (

3 in.) thick mild steel plate when operating the plasma and gas-metal arcs simultaneously. This 8

compares to three passes when just the PAW process is used. One would expect the same joint to require two or three passes for the conventional GMAW process. When butt welding aluminum alloy sheet and plate in thicknesses ranging from 4.6 to 6.1 mm (0.18 to 0.24 in.), singlepass welds can be made at travel speeds that are more than twice as fast as when the GMAW process is used alone (Table 2). Lower porosity levels are achieved with the plasma-MIG process because of the cleaning action of the plasma arc. When using mild steel and stainless steel, the plasma-MIG process can make butt welds comparable to those made with the GMAW process alone at almost twice the travel speed (Table 2).

TABLE 2 COMPARISON OF WELDING SPEEDS FOR GMAW AND PLASMA-MIG WELDING OF ALUMINUM AND STEEL PARAMETER

SHEET THICKNESS, MM (IN.) WIRE DIAMETER, MM (IN.) GMAW CURRENT, A PLASMA CURRENT, A WIRE FEED SPEED, M/MIN (IN./MIN) WELDING SPEED, M/MIN

MATERIAL ALMG4.5MN GMAW PLASMAMIG

PLASMAMIG

4.6 (0.18) 1.6 (0.063) 280-300 ... 8.1 (320)

4.6 (0.18) 1.6 (0.063)

5.1 (0.20) 1.6 (0.063)

200 200 10.9 (430)

0.4 (16)

1.1 (43)

STEEL GMAW

PLASMAMIG

280 200 10.9 (430)

4.1 (0.16) 1.2 (0.047) 250 ... 5.1 (200)

4.1 (0.16) 1.2 (0.047) 280 200 11.4 (450)

0.97 (38)

0.51 (20)

0.97(38)

(IN./MIN) Source: Ref 2

The hardfacing of dredging equipment with martensitic steel was made more productive by using the plasma-MIG process, rather than the conventional GMAW process. Higher deposition rates were achieved, resulting in reduced labor costs for the semiautomatic welding operation. Deposition rates of up to 20 kg/h (44 lb/h) can be achieved by plasmaMIG surfacing of mild steel with stainless steel flux-cored wires (Table 3).

TABLE 3 WELD SURFACING PARAMETERS FOR STAINLESS STEEL DEPOSITS ON MILD STEEL USING PLASMA-MIG WITH FLUX-CORED WIRES

UNDILUTE D WELD METAL PZ 6400 (347) SURFACIN G PZ 6410 (308) SURFACIN G PZ 6415 (309) SURFACIN G

MIG WELDIN G CURREN T, A

MIG WELDIN G VOLTAG E, V

PLASMA WELDIN G CURREN T, A

PLASMA WELDIN G VOLTAG E, V

WELDING SPEED

DEPOSITIO N RATE

INNER WIRE STICKOUT

mm/ mi

in./mi n

kg/ h

lb/h

m m

in.

300

39

150

44

110

4.4

10

22

35

200

31

150

40

80

3.2

6.6 14. 6

300

44

150

46

140

5.6

13

400

50

150

51

190

7.6

20

DILUTIO N, %

FERRIT E NO.

1. 4

...

15, 12, 13

55

2. 2

8.6

12

28. 7

55

2. 2

13.8

5.5

44

55

2. 2

9.7

6.5

Source: Ref 3

References cited in this section

2. J.D. SWART, PLASMA-MIG BOOSTS TANK TRAILER OUTPUT, WELD. DES. FABR., FEB 1983, P 54-55, 59 3. F. EICHHORN AND E. VAN GAEVER, ADVANTAGEOUS SURFACING WITH PLASMA-MIG USING CORED WIRES, PROC. 1ST INT. CONF. SURFACE ENGINEERING, THE WELDING INSTITUTE, 1986, P 99-110 Plasma-MIG Welding Ian D. Harris, Edison Welding Institute

Personnel Skill Level and Training. The plasma-MIG process can be used for semiautomatic, mechanized, or automated operation. It requires a skill level comparable to that of the conventional GMAW process, although additional parameters

need to be set for the plasma arc. Most welding and surfacing activities are carried out with mechanized equipment. When welding parameters are set, operation is similar to that of the GMAW process. Health and safety issues are similar to those of other arc welding processes. Electric shock, eye protection, burns, ultraviolet radiation, and fume exposure are typical concerns. The plasma-MIG process generates more radiant heat than the conventional GMAW process. Plasma-MIG Welding Ian D. Harris, Edison Welding Institute

References

1. W.G. ESSERS, A.C.H.G. LIEFKENS, AND G.W. TICHELAAR, PLASMA-MIG WELDING, PROC. CONF. ADVANCES IN WELDING PROCESSES, THE WELDING INSTITUTE, 1971, P 216-219 2. J.D. SWART, PLASMA-MIG BOOSTS TANK TRAILER OUTPUT, WELD. DES. FABR., FEB 1983, P 54-55, 59 3. F. EICHHORN AND E. VAN GAEVER, ADVANTAGEOUS SURFACING WITH PLASMA-MIG USING CORED WIRES, PROC. 1ST INT. CONF. SURFACE ENGINEERING, THE WELDING INSTITUTE, 1986, P 99-110 Resistance Spot Welding Neville T. Williams, British Steel

Introduction RESISTANCE SPOT WELDING (RSW) is a process in which faying surfaces are joined in one or more spots by the heat generated by resistance to the flow of electric current through workpieces that are held together under force by electrodes. The contacting surfaces in the region of current concentration are heated by a short-time pulse of low-voltage, highamperage current to form a fused nugget of weld metal. When the flow of current ceases, the electrode force is maintained while the weld metal rapidly cools and solidifies. The electrodes are retracted after each weld, which usually is completed in a fraction of a second. Spot welding is the most widely used joining technique for the assembly of sheet metal products such as automotive body-in-white assemblies, domestic appliances, furniture, building products, enclosures and, to a limited extent, aircraft components. Many assemblies of two or more sheet metal stampings that do not require gas-tight or liquid-tight joints can be more economically joined by high-speed RSW than by mechanical methods. Containers frequently are spot welded. The attachment of braces, brackets, pads, or clips to formed sheet-metal parts such as cases, covers, bases, or trays is another common application of RSW. Major advantages of spot welding include high operating speeds and suitability for automation or robotization and inclusion in high-production assembly lines together with other fabricating operations. With automatic control of current, timing, and electrode force, sound spot welds can be produced consistently at high production rates and low unit labor costs using semiskilled operators. Most metals can be resistance spot welded if the appropriate equipment is used coupled with suitable welding conditions. This is particularly true for thin sheet or strip steel products, whether uncoated or coated. Resistance Spot Welding

Neville T. Williams, British Steel

Equipment The equipment needed for RSW can be simple and inexpensive or complex and costly, depending on the degree of automation. Spot welding machines are composed of three principal elements: • • •

ELECTRICAL CIRCUIT, WHICH CONSISTS OF A WELDING TRANSFORMER, TAP SWITCH, AND A SECONDARY CIRCUIT CONTROL CIRCUIT, WHICH INITIATES AND TIMES THE DURATION OF CURRENT FLOW AND REGULATES THE WELDING CURRENT MECHANICAL SYSTEM, WHICH CONSISTS OF THE FRAME, FIXTURES, AND OTHER DEVICES THAT HOLD AND CLAMP THE WORKPIECE AND APPLY THE WELDING FORCE

Specifications for resistance welding equipment have been standardized by the Resistance Welder Manufacturers Association (RWMA), and specifications for controls are issued by the National Electrical Manufacturers Association (NEMA). Resistance Spot Welding Neville T. Williams, British Steel

Electrical Circuit The transformer used in a direct energy resistance delivers the power to the workpiece by changing the input highvoltage, low-amperage, alternating current (ac) in the primary winding to a low-voltage, high-amperage current in the secondary winding. This system forms the basis of pedestal, or gun-type, welding machines where the output of the secondary transformer is applied directly to the welding electrodes. More complex secondary circuits are used in multiwelders. To minimize the size and cost of multiwelders, transformers can be designed with two or more secondary circuits supplied from the same primary circuit. This design is achieved by using transformers in which the primary current is supplied to two or more separate secondary windings. Split or multiple secondary transformers are generally used to produce four or more welds at the same time. One controller is used for control of the timing and current supplied to the electrode, which means that it is not possible, under these circumstances, to achieve individual control or setting of the welding conditions for each pair of welding electrodes. Difficulties can arise in optimizing the necessary welding conditions to maximize the electrode life for each welding station under these conditions. In addition, the possibility exists for developing shunt currents, or an out-of-balance welding current, between each secondary circuit. Multiwelders make large use of single-welding and multiple-welding configurations based on direct, indirect, series, and push-pull secondary circuits (see Fig. 7 of the article "Procedure Development and Practice Considerations for Resistance Welding" in this Volume). Single-Weld Configurations. Most pedestal and portable gun-welding machines employ direct welding, where the

current flows directly through the welding electrodes and the workpiece to produce a single weld. In areas of difficult access, indirect welding may be necessary. Indirect welding is generally used when welding is carried out using either stationary or indexing welding guns and when access to the sheets being welded is limited to one side. In this case, the backing electrode generally consists of a large bar or platen, and the welding current flows from one side of the transformer secondary to the active electrode, through the welded area, and then along the platen/bar electrode or connecting jumper cable to the dummy electrode, which is connected to the other leg of the transformer secondary. Higher welding currents are generally necessary when using dummy welding guns, which can lead to a shorter electrode life. The extent to which larger currents are required depends largely on the distance between the welding electrode and the dummy return electrode. In a multiwelder, this distance can be either fixed or variable, depending on whether a

stationary or indexing head is used. Both types of welding heads can lead to losses due to stray current paths and extensive shunting of the current through jigs and fixtures. These factors need to be considered in any design concept for a multiwelder station. Two-Weld Configurations. Because of the need to minimize the number of transformers and the overall size and cost of

a multiwelder, it is necessary to utilize configurations whereby two welds are made in one stroke of an electrode. Parallel, series, or push-pull welding configurations are used for this purpose. In parallel welding, two or more direct welds, usually closely spaced, are made simultaneously with all upper

electrodes connected to one side of the transformer, and all bottom electrodes connected to the other side of the transformer. Parallel welding has the advantage of lower shunt currents, thereby allowing shorter interweld spacings to be used. However, if two or more circuits are fed from a common transformer secondary winding, the current in each circuit will not be equal unless the impedances of each circuit are identical. This is often difficult to achieve in practice, and the electrode life obtained can, therefore, be different for each of the electrode pairs connected to the same transformer. Series and Push-Pull Welding. Two simultaneous welds can also be produced by means of series welding in which the

secondary circuit is connected to two contoured electrodes that contact the workpiece from the same side, and a current backup bar/platen is used on the reverse side. Push-pull welding is also employed whereby either one or two welds are made, depending on the electrode configuration. In push-pull RSW, two transformers are used to form the circuit between the welding electrodes; such a configuration has advantages, particularly where a deep-throated machine is required. Such machines ordinarily require a large secondary loop because of the need for conducting the welding current around the large working area. Using a push-pull configuration, where the upper set of transformers matches the lower set, eliminates the need for long cable loops to the electrodes, resulting in a lower secondary impedance. A longer electrode life can be obtained in push-pull welding compared to series welding. It should also be noted that current shunting between the electrodes occurs in both series and push-pull welding configurations, but to a lower extent in the latter. The magnitude of the shunt currents depends on interelectrode distance and the type of material being welded. In all cases, current shunting will lead to a lower electrode life when welding zinccoated steels due to the need to use higher welding currents to compensate for the shunt currents (this can also lead to additional electrode sticking problems) and to the rapid falloff in weld size that occurs under conditions where current compensation is insufficient or nonexistent. To some extent, the effects of both high secondary circuit impedance and current shunting can be overcome by using a constant-current facility fitted into the welding machine control circuit, but there is a limit to the effectiveness of these controls in a production situation. In addition, direct current (dc) welding systems can be used in which secondary inductance/reactance effects are minimized. However, the effectiveness of using dc depends on the method used to develop the dc waveform and on the extent of the superimposed ac ripple on the dc waveform. Secondary Impedance. The electrical characteristics of a spot welding machine are best defined in terms of their effect on the impedance of the secondary circuit. The latter is composed of two components: •

•

REACTANCE, WHICH IS DETERMINED PRIMARILY BY THE DEPTH AND HEIGHT OF THE THROAT OF THE MACHINE AND THE AMOUNT OF STEEL IN THE THROAT. ANOTHER CRITICAL FACTOR IS THE FREQUENCY OF THE APPLIED CURRENT. RESISTANCE, WHICH DEPENDS ON THE LENGTH AND CROSS SECTION OF THE CONDUCTORS. CONDUCTOR MATERIAL AND OPERATING TEMPERATURE NEED ALSO TO BE CONSIDERED.

Cables can be classified into three types: air-cooled jumper cables, water-cooled jumper cables, and kickless cables. In any application, it is necessary to ensure that the type and dimensions of cables are chosen to achieve the best balance between the total circuit resistance and the electrode life to be obtained with a multiple spot welder. Resistance Spot Welding Neville T. Williams, British Steel

Machine Construction Welding operations in highly automated production lines are based primarily on multiple spot welders and robotic cells. In addition, manual welding operations can be used to manufacture either subassemblies, which are fed into the main production/assembly lines, or, in many instances, finished products. These differing end uses require welding machines of varying designs and characteristics. RSW machines can be divided into three basic types: • • •

PEDESTAL-TYPE WELDING MACHINES PORTABLE WELDING GUNS MULTIPLE SPOT WELDING MACHINES INCORPORATING LIGHTWEIGHT GUN WELDING UNITS

Pedestal-type welding machines form the basis of most resistance spot or projection welding operations. Standard

machines are either of the rocker arm type, as determined by the rocker action of the moving upper electrode arm, or stationary, direct-acting machines. The latter outnumber the former because rocker arm machines are difficult to align due to the movement of the electrode in an arc motion. As the electrode wears, various skidding actions can be created between the upper and lower welding electrodes, depending on the extent of the electrode wear. This can change the point at which the electrodes meet relative to the fulcrum point of the moving arm. If the electrodes meet at a point above the fulcrum point, they have a tendency to slide outward from the machine. On the other hand, if the electrodes meet on a line below the fulcrum of the moving arm, the electrodes show a tendency to slide inward toward the machine. These relative movements can be different between the top and bottom electrodes, a consequence of which is a reduction in electrode life, particularly if welding coated steels. Other advantages of direct-acting machines include a lower electrode assembly mass, faster "follow up" during welding, and a degree of rigidity that is impossible to obtain with an ordinary rocker arm machine. Press-type direct-acting machines have an upper electrode and welding head that move vertically in a straight line. The welding head is guided in bearings or guideways of sufficient proportion to withstand the offset loads imposed upon them. All of these factors tend to increase the electrode life obtained in a particular welding cell. Portable welding guns are used in RSW when it is impractical or inconvenient to transport the work to the machine.

The portable welding gun consists of water-cooled electrode holders, an air or hydraulic actuating cylinder, hand grips, and an initiating switch. The gun usually is suspended from an adjustable balancing unit, and the welding force is supplied by an air cylinder. Because of the high secondary losses of portable machines, transformers used in these machines have a secondary voltage two to four times as great as the voltage of transformers used in stationary machines of equal rating. Welding current is transmitted between the transformer terminals and gun terminals through a secondary cable, usually of the low-impedance or kickless type. The reactance of this type of cable is near zero, which results in a high power factor and reduced kilovolt-ampere (kVA) demand. For safety reasons, the current initiating switch is usually operated at low voltage. Recent developments in encapsulated transformers and high-speed fail-safe devices enable the transformer to be integral with the welding gun. This results in lower secondary losses and is particularly preferred in robotic applications. Gun welders typical of portable manually operated and robotic welding cells can be classified into two basic types: • •

S-TYPE (SCISSORS) J-TYPE (DIRECT ACTION OR C/J TYPE)

Two important features that influence the performance of these types of welding guns are the rate of force buildup and the extent to which the arms of the gun deflect under load. These features characterize the welding gun in terms of rigidity and the mechanics involved in their operation. An S-type gun utilizes a lever action to transmit the force applied by the cylinder head to the welding electrodes. However, due to an unbalanced pivot point, the force measured at the electrodes is less than that applied at the cylinder. With J-type welding guns, the force applied by the pneumatic cylinder is transmitted directly to the welding electrode. The rigidity of the framework of the welding gun is characterized by the extent to which the electrode arms deflect on application of the electrode force. This is an important factor in determining weld quality and the amount of skidding that can occur between the two welding electrodes. It should be noted that extensive skidding can lead to elliptical welds and/or decreased electrode life.

Resistance Spot Welding Neville T. Williams, British Steel

Welding Electrodes Materials for RSW electrodes should have sufficiently high thermal and electrical conductivities and sufficiently low contact resistance to prevent burning of the workpiece surface or alloying at the electrode face. In addition, the electrode should have adequate strength to resist deformation at operating pressures and temperatures. Because the part of the electrode that contacts the workpiece becomes heated to high temperatures during welding, hardness and annealing temperatures must also be considered. Electrode materials for RSW have been classified by RWMA and in International Standards Organization (ISO) standard ISO 5182. When welding two sheets of thickness up to 3 mm (

1 in.) using truncated cone-type electrodes, the electrode tip diameter 8

should be determined from:

D2 = t

(EQ 1)

where d2 is the initial tip diameter (in mm) and t is the thickness (in mm) of the sheet in contact with the electrode. When using truncated cone electrodes, the initial or set-up weld diameter should be equal to the diameter of the electrode tip:

D = D2 = 5 t

(EQ 2)

where d is the weld diameter (in mm). Whenever welding two sheets of dissimilar thicknesses, the electrode dimensions and the required weld size should be specified appropriate to the thinner sheet thickness. In the case of three thicknesses, the second thinnest sheet should be used. When using domed or pointed electrodes, Eq 1 may not apply and the electrode dimensions will depend on accessibility and flange width. In this case, the electrode tip dimensions and welding conditions should be chosen to give an initial weld diameter as specified in Eq 2. Where a pad or mandrel is used as the second electrode, its surface must be maintained to match the profile of the work. During normal production, electrodes tend to mushroom leading to an increase in electrode tip size. According to trends in automotive practice and international standards, the diameter of at least one of the electrode tips should not normally be allowed to increase above a value that results in a reduction greater than 30% from the starting weld diameter during a production run. This is equivalent to an increase in the electrode diameter to 1.3d. When this diameter has been reached, the electrode should be replaced or redressed to its original size and contour. Electrode shapes have been standardized by RWMA. Figure 1 shows the six standard face or nose shapes, identified by

letters A through F. Electrodes with type A (pointed) tips are used in applications for which full diameter tips are too wide. Type D (eccentric, formerly called offset) faces are used in comers or close to upturned flanges. Special tools are available for dressing the electrode faces, either in or out of the welding machine.

FIG. 1 STANDARD-RESISTANCE SPOT WELDING ELECTRODE FACE AND NOSE CONFIGURATIONS

Electrode Coolant Parameters. Generally, it is recommended that the water flow to the electrodes should be a 1 minimum of 4 L/min (1.1 gal/min) for welding two uncoated steel sheets of thickness up to and including 3.0 mm ( in.). 8

Higher flow rates are recommended when welding coated steels. The internal water cooling feed tube should be arranged to ensure that the water impinges onto the back working face of the electrode. The distance between the back and the working face of the electrode should not exceed the values given in the appropriate ISO or RWMA standard. The inlet water temperature should not exceed 20 °C (70 °F), and the outlet temperature should not exceed 30 °C (85 °F). To maintain these temperature levels, the electrode cooling water supply should be independent of transformer and thyristor water cooling circuits. Separate water circuits should be used for both top and bottom electrodes. Resistance Spot Welding Neville T. Williams, British Steel

Weldability of Steel Uncoated Steels. Low-carbon steel can be satisfactorily resistance spot welded using a wide range of time, current, and electrode force parameters. Practices recommended for spot welding SAW 1010 steel have been issued by the RWMA (Ref 1) and the American Welding Society (Ref 2). These data can serve as starting points for the establishment of optimum settings but may require modification, depending on the type of welding machine, the dynamic properties of the welding machine, the pneumatic characteristics, the characteristics of the secondary circuit, the electrode shape, and the material. Sufficient squeeze time should be chosen to enable the electrode force to build up to its preset value. For lightweight gun welders with limited force capability, the electrode force values are reduced up to 30% for sheet

thicknesses greater than 1.6 mm (

1 in.). The welding current needs to be adjusted accordingly. When welding sheets of 16

dissimilar thickness, welding conditions may be based on the thinner sheet, or on the second thinnest sheet when welding three thicknesses. In the case of high-strength low-alloy (HSLA) steels, up to 20% higher electrode force may be necessary. Welding currents may be reduced up to 20%, depending on the type of high-strength steel being welded. Pulse welding schedules may be necessary at the higher carbon equivalent levels and at sheet thicknesses greater than 2.0 mm (

5 in.). The 64

welding conditions required may depend greatly on the analysis of the steel because of differences in the resistivity of the various types of high-strength steel. For example, carbon content has the greatest effect on weldability of steels; weld

hardness increases rapidly with small increases in carbon content. This high hardness causes nugget interfacial tears and nugget deterioration. To obtain acceptable weld performance, carbon content should be kept below 0.10% + 0.3t, where t is the sheet thickness in inches. For materials above this range, postweld tempering may be necessary. Low-carbon steels are generally considered to be spot weldable. These steels can be obtained from a number of suppliers with reasonable certainty of only minor variations in weldability. On the other hand, HSLA steels are sold on the basis of strength, with each steelmaker producing its own composition. Therefore, this variation in composition may cause variations in weldability, which the steel user must consider. Specific alloying elements or combinations of elements can impart desired nugget properties in low-carbon or HSLA steels. The effects of particular elements are discussed below. Phosphorus and Sulfur. Phosphorus and, to some extent, sulfur, are generally considered to promote nugget interface

tearing. When the total content of phosphorus, sulfur, and carbon exceeds a critical value, spot weld interfacial failure is observed. This interfacial tearing during peel testing results in reduced current ranges and minimal lobe curve areas. Titanium. Data on the effect of titanium on both hot-rolled and cold-rolled steel indicate that the spot weldability of these

steels is not as good as that of the HSLA steels containing niobium and/or vanadium. Increasing titanium content generally reduces maximum button diameter, shear and cross-weld tensile strength, and welding current range. Titanium content should not exceed 0.18%, and the use of oversized electrodes and higher force may be required when welding titanium-bearing steels. Nitrogen. In HSLA steels, nitrogen promotes interfacial nugget failure. Nitrogen appears to be more critical in unkilled

cold-rolled steel. However, the sensitivity can be reduced by decreasing the nitrogen content or by combining the nitrogen with aluminum as in aluminum-killed steels. Zinc-Coated Steels. The present trend in the automotive industry toward the use of larger amounts of zinc-coated steels

in automotive assemblies demands that certain strict guidelines regarding the selection of equipment, choice of welding schedules, and maintenance procedures are rigidly followed. Two basic types of alloy coating are available: • •

ZINC-IRON (BOTH HOT-DIP AND ELECTRODEPOSITED) ELECTRODEPOSITED ZINC-NICKEL (SUPPLIED WITH OR WITHOUT AN ORGANIC PRIMER TOP COAT)

The available welding ranges for zinc-iron and zinc-nickel coated steels are similar to those for uncoated mild steel although displaced toward slightly higher currents (Fig. 2). When these steels are coated, it is essential that the film thickness not exceed 1 to 1.5 μm (40 to 60 μin.) in order to facilitate breakthrough of the film to enable current flow between the welding electrodes at the low secondary voltages typical of RSW machines.

FIG. 2 PLOT OF WELD TIME VERSUS SECONDARY WELD CURRENT TO OBTAIN WELDABILITY LOBES FOR SELECTED 0.8 MM (0.03 IN.) THICK STEELS. ELECTRODE PARAMETERS: FORCE, 1.8 KN (0.20 TONF); TIP DIAMETER, 5.0 MM (0.20 IN.)

Welding conditions can be optimized easily when using pedestal-type welding machines; however, this may not be the case in body-in-white assembly where a combination of multiple spot welders and robotic cells are more commonly used. Such operations are susceptible to poor fit-up of parts, access problems, and electrode configurations that give rise to electrode skidding. Consequently, the welding parameters may need to be adjusted to compensate for such situations. In many instances, one welding gun is called on to weld a range of sheet thickness combinations or even a variety of material combinations in a particular part. Although a multi-program welding control can be used to adjust for the different welding currents necessary for the various combinations, the actual current values may not be optimum because, more often than not, a single compromise electrode force value has to be used. In these circumstances, the electrode life obtained may be lower than that derived using optimized welding conditions. For example, the electrode life obtained can depend greatly on the type of welding gun used (Fig. 3) or the angle of approach of the electrode. For multiple spot welders, it is essential that all welding stations are balanced electrically and that the same air pressure is supplied to each station. Electrode shape and configuration, water cooling arrangements, and electrode dressing schedules need to be optimized. The last item is most important because satisfactory weld quality can only be provided when the necessary electrode condition and shape are maintained over a production run.

FIG. 3 EFFECT OF RSW ELECTRODE GUN TYPE ON ELECTRODE LIFE FOR SELECTED STEELS AND ALLOYS

References cited in this section

1. RESISTANCE WELDING MANUAL, 4TH ED., RESISTANCE WELDER MANUFACTURERS ASSOCIATION, 1989 2. RECOMMENDED PRACTICES FOR AUTOMOTIVE WELD QUALITY--RESISTANCE SPOT WELDING, ANSI/AWS D8.7-88, AMERICAN WELDING SOCIETY, 1988 Resistance Spot Welding Neville T. Williams, British Steel

Weldability of Aluminum Alloys Most of the commercial aluminum alloys produced in sheet form can be spot welded provided that suitable welding equipment is used. Because of the low electrical resistance of aluminum alloys, considerably higher welding currents are required relative to welding mild steel. In addition, the narrow plastic range between softening and melting means that welding forces, time, and current need to be closely controlled. For consistent weld quality, it is essential that the tenacious surface oxide films are removed by mechanical or chemical techniques prior to welding. To minimize the likelihood of cracking or porosity in the weld, careful control of the electrode force is necessary. If the force is applied too soon, the welding current may be insufficient due to a premature loss of contact resistance. Thus, correct initiation of the force relative to the current is essential. Dual force schedules are frequently used; therefore, a lowinertia welding head can be beneficial because it allows a more effective buildup of force to its final level. Because of the higher secondary currents required when welding aluminum alloys, direct current secondary rectified or frequency converter welding machines are favored, particularly if high-quality welds are required (such as in aircraft applications).

Resistance Spot Welding Neville T. Williams, British Steel

References

1. RESISTANCE WELDING MANUAL, 4TH ED., RESISTANCE WELDER MANUFACTURERS ASSOCIATION, 1989 2. RECOMMENDED PRACTICES FOR AUTOMOTIVE WELD QUALITY--RESISTANCE SPOT WELDING, ANSI/AWS D8.7-88, AMERICAN WELDING SOCIETY, 1988 Projection Welding Jerry E. Gould, Edison Welding Institute

Introduction PROJECTION WELDING (PW) is a variation of resistance welding in which current flow is concentrated at the point of contact with a local geometric extension of one (or both) of the parts being welded. These extensions, or projections, are used to concentrate heat generation at the point of contact. The process typically uses lower currents, lower forces, and shorter welding times than does a similar application without the projections. Projection welding is often used in the most difficult resistance-welding applications because a number of welds can be made at one time, which speeds up the manufacturing process. PW applications are generally categorized as being either embossed-projection welding or solid-projection welding. These variations are shown in Fig. 1 and 2. Embossed-projection welding is generally a sheet-to-sheet joining process, in which a projection is stamped onto one of the sheets to be joined. Then, resistance welding is conducted on a stack of sheets. Heat initially concentrates at the contact point and in the walls of the projection during resistance welding. Early in the process, the projection itself collapses back into the original sheet. However, the initial heating raises the local resistivity of the joint area, allowing continued resistance heating at this location. Weld development then proceeds in the conventional manner, by forming a fused weld nugget.

FIG. 1 TYPICAL STACKED CONFIGURATION FOR EMBOSSED-PROJECTION WELDING OF SHEET

FIG. 2 TYPICAL CONFIGURATION FOR SOLID-PROJECTION WELDING

Solid-projection welding requires that the projection be forged onto one of the two components. Then, during resistance welding, the contact point and the projection itself experience preferential heating. In this case, the projection cannot simply collapse, as it does in embossed-projection welding. Rather, the projection collapses by penetrating the opposing material and by extrusion to the periphery. When compared with embossed-projection welding, the resulting joints are solid-state, rather than fusion, welds. The actual joints are caused by a combination of material forging and diffusion bonding, much like they are in resistance butt and flash butt welding. Projection Welding Jerry E. Gould, Edison Welding Institute

Applications Examples of projection welding are shown in Fig. 3. These applications, which range from sheet-to-sheet joints, to crosswire welds, to annular attachments, to nut welds, to weld screws, include both embossed- and solid-projection types of welding. Specific welding examples, as well as material effects, are described below.

FIG. 3 DIFFERENT PROJECTION-WELDING CONFIGURATIONS

Embossed Projections Typical embossed-projection weld configurations are shown in Fig. 3(a), (b), (d), and (h). The embossed-projection technique is primarily used for sheet-to-sheet applications. It does involve the additional expense of stamping the projections and ensuring that the electrodes are properly located over the projections. However, the ability of the projection to create highly localized joints allows this technique to extend the applications of resistance welding. Most notably, projection welding widely increases the allowable thickness mismatch associated with resistance welding. Generally, resistance welding is considered adaptable to thickness variations of less than 3:1. However, with the addition of the projection, there is no limit to the difference in joinable thicknesses. Similarly, embossed-projection welding is highly adaptable to the joining of multiple-sheet stacks. The approach is similar in that the projections are used to localize heat at critical interfaces in the overall joint. Another more recent use of embossed-projection welding is in applications where external marking must be minimized. This is a concern when a primer has been applied to one side of one of the sheets being joined. Because it is very effective at localizing heat at the interface, and because it uses shorter welding times, the PW process is more adiabatic and there is less residual heat that can damage outside surfaces. However, surface marking still does occur.

Solid Projections A considerably wider range of solid-projection welding processes is commonly used in production applications. Annular projection welding, like embossed-projection welding, is commonly used to provide a highly localized joint and to minimize thermal damage to other parts of the structure. Common variations of solid-projection welding are described below. Annular Projection Welds. One of the most common applications of solid-projection welding is to attach either tubular

components or members with circular bases to flat or curved substrates. This is commonly accomplished by annular projection welding, in which the projection is machined onto the circular base or tubular section. Then, resistance welding is conducted to set the projection into the substrate. A typical weld geometry is shown in Fig. 3(g). Annular projection welding generally provides high-integrity joints and can be used for leak-tight applications. Cross-wire welding is a variation of solid-projection welding in which the projection is formed by the contact point of

two crossing wires. The geometry for cross-wire welding is shown in Fig. 3(n). Upon resistance welding, heat is maximized at the location of the wire-wire contact, and the parts are subsequently forged together. Depending on the application, either highly localized joints (minimal forging) or heavily forged joints can be made. In some applications, cross-wire welding is conducted at rates up to 30 welds/s. Weld nuts represent an application of solid-projection welding in which the projections themselves occur as actual extensions of material from the nut. Typically, weld nuts contain 3, 4, or 6 of these projections, or feet, spaced around the nut periphery. During welding, all projections are attached simultaneously. Examples of weld nut geometries are shown in Fig. 3(k) and 3(m). Edge-to-sheet welds are typically a cross between embossed- and solid-projection welding. They are used to attach the

end of a sheet to the flat face of an opposing sheet (Fig. 3f). The projections for welding are generally stamped into the face of the attaching sheet. During welding, the projection does collapse locally back into the base sheet. However, bonding is strictly solid state, similar to that of other solid-projection welding processes. Material Effects Solid-projection welds are essentially strain-assisted diffusion bonds. Because the projections typically collapse at very high temperatures (generally, within several hundred °C of the melting point), diffusion bonding can occur within the very limited available time (usually, less than 1 s). Not surprisingly, some of the material-related factors that affect diffusion bonding also affect solid-projection welding. The most notable of these is the ability of the metal to dissolve its own oxide. As a result, materials that do not energetically favor the solid solubility of oxygen, rather than the formation of the oxide at aluminum forging temperatures, will be relatively difficult to projection weld. The strength-temperature relationship also affects projection weldability. Materials that maintain their strengths at relatively high temperatures permit substantial heating to occur before projection collapse. This heat then becomes available to promote diffusion after projection collapse. Premature collapse results in lower temperatures in which diffusion can occur and in reduced current density, which prohibits further resistance heating. Bulk resistivity also plays a role in projection welding, but to a lesser degree. Increased bulk resistivity can reduce the effectiveness of the projection as a current concentrator. With increasing bulk resistivity, the tendency is for delocalized heating and general, rather than local, collapse of the projection. As a result, high-resistivity materials are more difficult to projection weld. Mild steels and low-alloy, nickel-base alloys are ideal materials for projection welding, because they readily dissolve their own oxides and have adequate strength-temperature and resistivity properties. Stainless steels and higher-alloy nickelbase materials become slightly more difficult to weld, because of the formation of more-stable chromium and aluminum oxides, increased high-temperature properties, and higher resistivities. Projection welding is commonly applied to copper and copper-base alloys. In many applications, projections are virtually required for resistance welding, because of the high conductivity of these materials.

On the other hand, aluminum and aluminum-base alloys are very difficult to projection weld. The aluminum oxide is so tenacious that solid-projection welding, in particular, is nearly impossible. In addition, it is very difficult to localize heat, because aluminum alloys soften at such low temperatures. Conventional titanium alloys are also relatively difficult to projection weld. Although titanium readily dissolves its own oxide, its high resistivity and low forging temperature generally cause premature collapse of the projection. Projection Welding Jerry E. Gould, Edison Welding Institute

Equipment Resistance projection welding generally utilizes slightly modified resistance-welding equipment. Because the nature of projection welding is to direct current flow to the desired weld locations, there is considerably more flexibility in tooling. A specific example is spot welding that requires electrodes of specific sizes and shapes in order to localize current for welding. Projection welding is typically done with large, flat electrodes. In most applications, tooling is simply shaped to match the contour of the part in the contacting location. Given the current-locating characteristic of projections, multiple projection welds can be made simultaneously. Projection welding does require high compliance or the rapid response of the loading system. This is necessary because projection welding requires the collapse of a projection, which, in turn, requires some motion of the welding head. If the follow-up of the welding head is insufficient, then a local loss of force will occur, potentially causing catastrophic overheating of the contacting surfaces. Therefore, PW systems typically utilize "fast follow-up" heads (Fig. 4), which consist of relatively low inertia (mass) head components and mechanisms to maintain a nominally constant force. The head shown in Fig. 4 uses a relatively low volume diaphragm assembly and a spring to maintain this force on the components. Most systems, however, use either these low-volume diaphragm assemblies or springs to accomplish fast follow-up.

FIG. 4 TYPICAL FAST FOLLOW-UP (LOW INERTIA) HEAD FOR PROJECTION WELDING. (A) IN OPEN POSITION. (B) IN POSITION FOR SQUEEZING AND HEATING THE PROJECTION. (C) AT INSTANT OF PROJECTION COLLAPSE AND START OF NUGGET FORMATION

Projection Welding Jerry E. Gould, Edison Welding Institute

Personnel Like all resistance welding processes, PW processes utilize basically automated equipment, which means that the expertise level to operate the equipment is relatively low. Generally, the skill levels and the health and safety issues are identical to those of conventional resistance spot welding. An area in which expertise is beneficial to hardware function and on-stream weld quality is operator attention to detail. Resistance welding is subject to variations in setup conditions, and projection welding may be extraordinarily sensitive to setup variations in situations where the placement of projections is critical. With limited expertise, operators or supervisory staff can visually inspect the characteristics of the process and the welds themselves to identify improper setups, as well as maintenance needs. These quality assurance efforts, which fall under the auspices of total quality management or total quality joining, can be established with a minimum of on-line training and experience. Projection Welding Jerry E. Gould, Edison Welding Institute

Process Requirements The types of hardware required for projection welding are analogous to those required for other resistance-welding processes. The embossed-projection welding of heavy-gage, intermediate-gage, and thin-gage sheet, as well as joints of

dissimilar thicknesses, is described below. Heavy-Gage Sheet Steels. The projection welding of heavy-gage steels, which involves an embossed-projection

welding process, has many of the characteristics of solid-projection welding particularly in its early stages, because of the thickness of the sheet. In addition to defining the geometry of the projection, projection designs should provide an annular relief for projection material that is forged to the side during welding. Projection and die geometries for steels that range from 3.12 to 6.22 mm (0.123 to 0.245 in.) are shown in Table 1. The process requirements for forming these welds are given in Table 2. All welding schedules described in this article are single-pulse welding schedules.

TABLE 1 PROJECTION AND DIE GEOMETRIES FOR WELDING A RANGE OF HEAVY-GAGE STEELS Data are for tool steel hardened to 50-52 HRC.

USS GAG E NO.

11 10 9 8 7 6 5 4 3

MATERIAL THICKNES S (T )

PROJECTION HEIGHT (H DIAMETER ), (D ), ±5% ±2%

mm

in.

mm

in.

mm

in.

mm

in.

mm

in.

mm

in.

DIE HOLE DIAMETER (D), ±0.1 MM (±0.005 IN.) mm in.

3.1 2 3.4 2 3.8 8 4.1 7 4.5 4 4.9 5 5.3 3 5.7 2 6.2 2

0.12 3 0.13 5 0.15 3 0.16 4 0.17 9 0.19 5 0.21 0 0.22 5 0.24 5

1.4 7 1.5 7 1.6 3 1.7 3 2.0 3 2.1 3 2.2 8 2.5 4 2.8 4

0.05 8 0.06 2 0.06 4 0.06 8 0.08 0 0.08 4 0.09 0 0.10 0 0.11 2

6.8 6 7.6 2 8.3 8 9.1 4 9.9 0 10. 4 11. 2 11. 9 13. 5

0.27 0 0.30 0 0.33 0 0.36 0 0.39 0 0.41 0 0.44 0 0.47 0 0.53 0

4.9 8 5.4 6 5.9 7 6.3 0 6.9 6 7.2 6 7.7 5 8.2 6 9.2 7

0.19 6 0.21 5 0.23 5 0.24 8 0.27 4 0.28 6 0.30 5 0.32 5 0.36 5

2.3 9 2.7 7 3.1 8 3.5 8 3.9 6 3.9 6 4.7 5 4.7 5 4.7 5

0.09 4 0.10 9 0.12 5 0.14 1 0.15 6 0.15 6 0.18 7 0.18 7 0.18 7

1.9 1 2.0 6 2.1 6 2.3 1 2.6 4 2.8 2 3.0 5 3.3 5 3.7 6

0.07 5 0.08 1 0.08 5 0.09 1 0.10 4 0.11 1 0.12 0 0.13 2 0.14 6

5.6 1 6.3 5 6.8 6 7.5 4 8.3 3 8.5 9 9.0 9 9.3 5 10. 3

Source: Ref 1

RADIUS (S )

PUNCH RADIUS (R ), ±0.1 MM (±0.005 IN.)

HEIGHT (P), ±2%

0.22 1 0.25 0 0.27 0 0.29 7 0.32 8 0.33 8 0.35 8 0.36 8 0.40 6

RECESS RADIUS (R), R = S/3

RECESS HEIGHT (H), H = H/3

mm

in.

mm

in.

1.6 5 1.8 3 1.9 8 2.1 1 2.3 1 2.4 1 2.5 7 2.7 4 3.0 7

0.06 5 0.07 2 0.07 8 0.08 3 0.09 1 0.09 5 0.10 1 0.10 8 0.12 1

0.48 3 0.50 8 0.53 3 0.58 4 0.68 6 0.71 1 0.76 2 0.83 8 0.94 0

0.01 9 0.02 0 0.02 1 0.02 3 0.02 7 0.02 8 0.03 0 0.03 3 0.03 7

TABLE 2 PROCESS REQUIREMENTS FOR PROJECTION WELDING OF A RANGE OF HEAVY-GAGE LOW-CARBON STEELS USS GAGE NO.

THICKNESS (A)(B)

mm

PROJECTION SIZE

DIAMETER HEIGHT

in.

mm

in.

mm

in.

MINIMUM MINIMUM SPACING, CENTERLINE TO CENTERLINE mm in.

SCHEDULE A: WELDING NORMAL-SIZED WELDS 9 3.89 0.153 8.38 0.330 1.57 0.062 44.5 8 4.17 0.164 8.89 0.350 1.73 0.068 45.7 7 4.55 0.179 9.91 0.390 2.03 0.080 48.3 6 4.95 0.195 1.04 0.410 2.13 0.084 50.8 5 5.33 0.210 1.12 0.440 2.34 0.092 53.3 4 5.72 0.225 1.19 0.470 2.54 0.100 33.0 3 6.22 0.245 1.35 0.530 2.84 0.112 63.5 SCHEDULE B: WELDING SMALL-SIZED WELDS 9 3.89 0.153 6.86 0.270 1.47 0.058 40.6 8 4.17 0.164 7.37 0.290 1.57 0.062 41.9 7 4.55 0.179 7.87 0.310 1.70 0.067 43.2 6 4.95 0.195 8.38 0.330 1.83 0.072 44.5 5 5.33 0.210 8.89 0.350 1.96 0.077 45.7 4 5.72 0.225 9.40 0.370 2.08 0.082 48.3 3 6.22 0.245 9.91 0.390 2.24 0.088 53.3

MINIMUM CONTACT OVERLAP

ELECTRODE FORCE WELD

ELECTRODE FORCE FORGE

UPSLOPE TIME(D), CYCLES

WELDTIME(D), CYCLES

mm

in.

mn

lbf

mn

lbf

1.75 1.80 1.90 2.00 2.10 1.30 2.50

22.9 24.1 25.4 26.7 29.2 30.5 33.0

0.9 0.95 1.0 1.05 1.15 1.20 1.30

8.9 10.2 11.7 13.0 14.2 16.1 17.4

2000 2300 2630 2930 3180 3610 3900

17.8 20.5 23.4 26.1 28.3 32.1 34.7

4000 4600 5260 5860 6360 7220 7800

15 15 20 20 25 25 30

60 70 82 98 112 126 145

1.60 1.65 1.70 1.75 1.80 1.90 2.10

19.1 20.3 21.6 22.9 24.1 25.4 27.9

0.75 0.80 0.85 0.90 0.95 1.00 1.10

6.2 6.3 6.7 7.1 7.7 8.3 9.3

1400 1425 1500 1600 1730 1870 2100

12.5 12.7 13.4 14.2 15.4 16.6 18.7

2800 2850 3000 3200 3460 3740 4200

15 15 20 20 25 25 30

60 70 82 98 112 126 145

WELDING CURRENT(C), A

TENSILESHEAR STRENGTH(F)

mpa

ksi

15,400 16,100 17,400 18,800 20,200 21,500 23,300

52 56 66 78 86 103 119

7.5 8.1 9.5 11.3 12.5 15.0 17.3

11,100 11,800 12,800 13,900 14,900 16,000 17,300

35 38 45 53 59 72 83

5.1 5.5 6.5 7.7 8.5 10.4 12.0

Source: Ref 1

(A) LOW-CARBON STEEL, SAE 1005-1010, 290-380 MPA (42-55 KSI) ULTIMATE TENSILE STRENGTH. (B) SURFACE OF STEEL MAY BE OILED LIGHTLY BUT FREE FROM GREASE, SCALE, AND DIRT. (C) ON SINGLE FORCE WELDS USE ONLY WELD FORCE AS ELECTRODE FORCE. ELECTRODE FORCE CONTAINS NO FACTOR TO FURTHER FORM POORLY MADE PARTS. (D) BASED ON 60 HZ. (E) STARTING VALUES SHOWN ARE BASED ON EXPERIENCE OF RWMA MEMBER COMPANIES. (F) TENSILE-SHEAR STRENGTH PER PROTECTION DEPENDS ON THE JOINT DESIGN.

Slope control is recommended to prevent preflashing of the projection upon initiation of the welding current. Forge forces are also recommended, because weld porosity is often a concern with heavy-gage steel. Electrode or die wear may also be a concern, as it is in the spot welding of heavy-section steels. In such applications, pulsation welding schedules might be recommended. Intermediate-Gage Sheet Steels. The projection welding of steels ranging in gage from 0.56 to 3.43 mm (0.022 to 0.135 in.), using single-point projections with single-impulse welding schedules, is well established. Projection-stamping die designs are given in Table 3, whereas process requirements are given in Table 4. Lower forces and shorter welding times are required, when compared with the conventional spot welding of these gages.

TABLE 3 PROJECTION AND DIE GEOMETRIES FOR WELDING A RANGE OF INTERMEDIATE-GAGE STEELS TO MAKE SPHERICAL PROJECTIONS Data are for tool steel hardened to 50-52 HRC.

USS GAGE NO.

25-21 20-19 18-17 16-15 14 13 12 11 10

MATERIAL THICKNESS (T )

PROJECTION HEIGHT DIAMETER (H ), (D ), ±5% ±2%

mm 0.560.86 0.911.1 1.21.4 1.51.7 1.9 2.3 2.72 3.12 3.43

mm in. mm in. mm in. mm in. mm in. mm in. 0.64 0.025 2.3 0.090 9.53 0.375 0.79 0.031 1.93 0.076 2.29 0.090

in. 0.0220.034 0.0360.043 0.0490.054 0.0610.067 0.077 0.092 0.107 0.123 0.135

PUNCH DIAMETER (A)

POINT RADIUS (R ), ±0.05 MM (±0.002 IN.)

DIE HOLE DIAMETER (B ), ±0.1 MM (±0.005 IN.)

CHAMBER DIAMETER (D)

0.89 0.035 2.8

0.110 9.53 0.375 1.19 0.047 2.26 0.089 2.79 0.110

0.97 0.038 3.6

0.140 9.53 0.375 1.19 0.047 2.64 0.104 3.56 0.140

1.1

0.042 3.8

0.150 9.53 0.375 1.57 0.062 3.05 0.120 3.81 0.150

1.2 1.3 1.4 1.5 1.6

0.048 0.050 0.055 0.058 0.062

0.180 0.210 0.240 0.270 0.300

4.6 5.3 6.1 6.9 7.6

9.53 12.7 12.7 12.7 12.7

0.375 0.500 0.500 0.500 0.500

1.57 1.98 1.98 2.39 2.77

0.062 0.078 0.078 0.094 0.109

3.66 4.37 4.98 5.61 6.35

0.144 0.172 0.196 0.221 0.250

4.57 5.33 6.10 6.86 7.62

0.180 0.210 0.240 0.270 0.300

Source: Ref 1

TABLE 4 PROCESS REQUIREMENTS FOR PROJECTION WELDING OF A RANGE OF INTERMEDIATEGAGE LOW-CARBON STEELS Data are for SAE 1005-1010 steels with ultimate tensile strengths ranging from 290-380 MPa (42-55 ksi). Surface of steel may be oiled lightly but must be free from grease, scale, and dirt.

CONFIGURATION(A)

A B C D E F G H I J K CONFIGURATION(A)

USS GAGE NO.

25 23 21 19 18 16 14 13 12 11 10

MATERIAL THICKNESS

PROJECTION SIZE

DIAMETER HEIGHT

MINIMUM SPACING CENTERLINE TO CENTERLINE

MINIMUM CONTACT OVERLAP

mm 0.559 0.711 0.864 1.09 1.24 1.55 1.96 2.34 2.72 3.12 3.43

mm 2.29 2.29 2.79 2.79 3.56 3.81 4.57 5.33 6.10 6.86 7.62

mm 9.65 9.65 12.7 12.7 19.1 19.1 22.4 26.9 31.8 38.1 41.4

mm 6.35 6.35 9.65 9.65 12.7 12.7 12.7 15.7 19.1 20.6 22.4

WELD TIME(B), CYCLES

in. 0.022 0.028 0.034 0.043 0.049 0.061 0.077 0.092 0.107 0.123 0.135

in. 0.090 0.090 0.110 0.110 0.140 0.150 0.180 0.210 0.240 0.270 0.300

ELECTRODE FORCE(C)

mm 0.635 0.635 0.889 0.889 0.965 1.07 1.22 1.27 1.40 1.47 1.57

in. 0.025 0.025 0.035 0.035 0.038 0.042 0.048 0.050 0.055 0.058 0.062

WELDING CURRENT(D), A

in. 0.38 0.38 0.50 0.50 0.75 0.75 0.88 1.06 1.25 1.50 1.63

TENSILE-SHEAR STRENGTH(E)

kn lbf mpa ksi (F) WELDING SCHEDULE A (FOR SINGLE PROJECTION) A 3 0.67 150 4,400 2.6 0.37 B 3 0.87 195 5,500 3.4 0.50 C 3 1.1 240 6,600 4.8 0.70 D 5 1.5 330 8,000 7.3 1.06 E 8 1.8 400 8,800 9.0 1.30 F 10 2.4 550 10,300 12.4 1.80 G 14 3.6 800 11,850 16.7 2.43 H 16 4.5 1020 13,150 22.4 3.25 I 19 5.6 1250 14,100 26.5 3.85 J 22 6.7 1500 14,850 33.1 4.80 K 24 7.3 1650 15,300 37.9 5.50 WELDING SCHEDULE B (FOR 1-3 PROJECTIONS), EACH PROJECTION(G) A 6 0.67 150 3,850 2.3 0.33 B 6 0.67 150 4,450 3.0 0.43 C 6 0.67 150 5,100 3.7 0.53 D 10 0.93 210 6,000 6.1 0.88 E 16 1.2 270 6,500 7.6 1.10 F 20 1.6 365 7,650 11.0 1.58 G 28 2.4 530 8,850 14.8 2.15 H 32 3.0 680 9,750 19.3 2.80 I 38 3.7 830 10,600 23.8 3.45 J 45 4.5 1000 11,300 29.0 4.20 K 48 4.9 1100 11,850 33.4 4.85 WELDING SCHEDULE C (FOR [GES]3 PROJECTIONS), EACH PROJECTION(H)

in. 0.25 0.25 0.38 0.38 0.50 0.50 0.50 0.62 0.75 0.81 0.88

B C D E F G H I J K

8 11 15 19 25 34 42 50 60 66

0.45 0.56 0.71 0.98 1.5 2.1 2.7 3.3 4.0 4.5

100 125 160 220 330 470 610 740 900 1000

3300 3800 4300 4600 5400 6400 7200 8300 9200 9900

2.3 2.9 5.0 6.1 8.5 12.1 16.0 20.0 24.8 29.3

0.34 0.43 0.72 0.88 1.23 1.75 2.33 2.90 3.60 4.25

(A) CONFIGURATION DESIGNATIONS ARE FOR THE PURPOSES OF THIS TABLE ONLY. (B) BASED ON 60 HZ. (C) ELECTRODE FORCE CONTAINS NO FACTOR TO FURTHER FORM POORLY MADE PARTS. (D) STARTING VALUES SHOWN ARE BASED ON EXPERIENCE OF RWMA MEMBER COMPANIES. (E) TENSILE-SHEAR STRENGTH PER PROJECTION DEPENDS ON THE JOINT DESIGN. (F) SCHEDULE A IS USABLE FOR WELDING MORE THAN ONE PROJECTION IF CURRENT IS DECREASED, BUT EXCESSIVE WELD EXPULSION MAY RESULT AND POWER DEMAND WILL BE GREATER THAN THAT FOR SCHEDULE B OR C. (G) SCHEDULE B IS USABLE FOR WELDING MORE THAN THREE PROJECTIONS, BUT SOME WELD EXPULSION MAY RESULT, AND POWER DEMAND WILL BE GREATER THAN THAT FOR SCHEDULE C. (H) SCHEDULE C IS USABLE FOR WELDING LESS THAN THREE PROJECTIONS WITH WELD CURRENT INCREASED APPROXIMATELY 15% AND POSSIBLE OBJECTIONABLE FINAL SHEET SEPARATION. Thin-Gage Sheet. The projection welding of thin-gage steel sheet (75

1-3 >3

1 50

A parameter of particular interest to induction heating is the penetration depth of heating (also called the skin depth), and this is denoted by the symbol δ. The skin depth is defined as the position in the solid away from the surface at which the magnitude of the magnetic field falls to approximately 37% of its value at the part surface. The skin depth is given by: δ =

ρ πµ f

(EQ 2)

where ρ is the resistivity of the part, ρis the absolute magnetic permeability, and f is the frequency. The absolute permeability is equal to the product μrμ0, where μr is the relative permeability of the material (e.g., pure iron = 18,000 and nonmagnetic copper = 1) and μ0 is the permeability of "free space," which equals 1.257 × 10-6 V · s/m · A (a universal constant). Skin depth increases as the material resistivity increases and as the frequency decreases. For magnetic materials at temperatures below the Curie point, the value of μr is not a constant; it is dependent on the magnetic field intensity, H0. For example, the value of μr for a low-carbon 1044 steel is 734,000/ H 00.92 .

Reference cited in this section

1. P. SIMPSON, INDUCTION HEATING, MCGRAW-HILL, 1960

Reference cited in this section

1. P. SIMPSON, INDUCTION HEATING, MCGRAW-HILL, 1960 Induction Soldering Paul T. Vianco, Sandia National Laboratories

Safety Concerns The safety hazards accompanying the use of induction heating include electrical shock from contact with the coils as well as burns caused by the heated workpiece. A third hazard is the danger of steam explosion caused by arcing between the coil and the part. Typically, the heating coils are hollow copper tubing; they are cooled by the internal flow of water or a water-base coolant. The vapor cloud accompanying the spatter of flux during soldering can cause an electric arc that may puncture the copper tubing if the induction coil is close to the workpiece. Arcing can also be caused by movement of the coil windings through forces generated by the magnetic field. Water contacting the hot part produces an immediate, violent steam explosion. To prevent this occurrence, adequate clearance should be maintained between the coil and the part (typical minimum spacing is 6.4 to 13 mm, or 0.25 to 0.50 in.). Moreover, trial heating runs, in which the current is slowly increased to assess coil movement and part-to-coil shorting, should be performed prior to product assembly. Induction Soldering Paul T. Vianco, Sandia National Laboratories

References

1. P. SIMPSON, INDUCTION HEATING, MCGRAW-HILL, 1960 2. F. LOBKOWICZ AND A. MELISSINOS, PHYSICS FOR SCIENTISTS AND ENGINEERS, SAUNDERS CO., 1975, P 355 Wave Soldering Paul T. Vianco, Sandia National Laboratories

Introduction WAVE SOLDERING is one of the primary techniques for mass assembly of printed wiring boards involving through holes, surface mount devices, or a combination of these two technologies. A schematic of the wave soldering process is shown in Fig. 1. A solder "fountain" or "wave" is created by a pump located at the bottom of the solder pot; suitable baffles are mounted in the pot to direct the flow of solder into the desired configuration. The printed wiring board is placed onto a conveyor, which brings it into contact with the wave surface. Note from Fig. 1 that the circuit boards travel along the surface of the solder; molten alloy does not flow on top of the board. As the printed circuit board passes on the wave, the solder wets the surface-mount package leads, terminations, and exposed metal surfaces in the circuit board, and also fills plated through holes. This technique can produce several thousand solder joints in a matter of minutes.

FIG. 1 SCHEMATIC OF THE WAVE SOLDERING PROCESS. THE THREE IMPORTANT PROCESS CONTROL REGIONS ARE THE ENTRY (A), THE INTERIOR (B), AND THE PEEL-BACK REGION (C).

The implementation of wave soldering for surface-mount technology requires that the devices be glued to the substrate prior to wave soldering. Moreover, the entire surface-mount package must be able to withstand contact with the flux and temperatures of the molten solder without a loss of reliability. Small-outline transistors and discrete components (chip resistors, capacitors, etc.) are readily attached by wave soldering; larger packages (e.g., leaded or leadless chip carriers) may be damaged by exposure to the harsh environments (flux and molten solder) as well as be more prone to solder defects (Ref 1). The wave soldering technique encompasses a sequence of processes, all of which are typically contained in the same apparatus. First, the substrate receives a coating of flux. Flux application methods include: • •

•

A WAVE TECHNIQUE SIMILAR TO THAT SHOWN IN FIG. 1 COATING THE CIRCUIT BOARD BY A FLUX FOAM CREATED BY PASSING AIR OR NITROGEN THROUGH A FLUX BATH TO GENERATE THE FOAM (OR FROTH) ON THE BATH SURFACE DIRECTLY SPRAYING THE FLUX ONTO THE BOARD SURFACE

After flux is applied, the substrate is passed through a preheating stage. Warming the board promotes activation of the flux, accelerates the evaporation of volatiles from the flux (which can cause voids and solder balls by spattering upon contact with the hot solder bath), and reduces thermal shock to the substrate and devices when it passes onto the solder wave. Then, the circuit board contacts the solder wave for the formation of the joints. After passing the wave, the board cools through natural heat loss or, more quickly, by the use of forced air. Acknowledgement This work was performed at Sandia National Laboratories and supported by the U.S. Department of Energy under contract DE-AC04-76DP00789.

Reference

1. R. PRASAD, SURFACE MOUNT TECHNOLOGY: PRINCIPLES AND PRACTICES, VAN NOSTRANDRHEINHOLD, 1989, P 426-428

Wave Soldering Paul T. Vianco, Sandia National Laboratories

Design Considerations and Process Parameters The successful assembly of a printed circuit board by the wave soldering technique relies on the proper process parameters and the layout of the board for the respective technology (through-hole, surface-mount, or mixed). The hole or via size (diameter) must permit adequate capillary filling by the molten solder and form the joint or to interconnect within the short time period (5 to 7 s) that the board contacts the wave. Holes with small diameters (1.0 mm, or 0.039 in.) and gaps (>0.40 mm, or 0.016 in.) decrease the capillary driving force of the solder to fill the volume and create the joint; in some cases, the solder simply drains from the hole under its own weight. In addition to insufficient solder in holes and vias, other defects in wave-soldered products include bridges, icicles, and skips. Bridges are defects in which a quantity of solder forms a short across two conductors (leads, lands, or traces). A piece of solder that does not short two conductors, but rather hangs from one surface, is an icicle. Skips refer to the failure of solder in the wave to reach the metal surface and form the joint. Design Considerations Design considerations, such as the orientation of through-hole leads and surface-mount devices with respect to the direction of travel past the wave, as well as the proximity of packages to 0one another, can dramatically affect yields (Ref 1, 2). Figure 2 is a diagram showing the preferred orientation of through-hole and surface-mount packages.

FIG. 2 SCHEMATIC OF PREFERRED DEVICE ORIENTATION FOR WAVE-SOLDERED ASSEMBLIES. SOIC, SMALLOUT-LINE INTEGRATED CIRCUIT (SURFACE-MOUNT); DIP, DUAL-INLINE PACKAGE (THROUGH-HOLE)

Through-Hole Technology. In the case of through-hole assemblies, the leads on the side of a through-hole 16-pin dual-

in-line package (DIP) should be oriented perpendicular to the travel direction of the conveyor to prevent bridging between them. Moreover, the bridging of leaded packages can be limited by maintaining the lead extension beneath the circuit board at less than 1.0 mm (0.039 in.). Leads that are too close to each other also increase the chances of solder bridges. Solder skips are less frequently observed on through-hole products (assuming adequate wettability of the metal surfaces) because the device package body is on the top surface, and therefore does not shadow joints from the wave.

Surface-Mount Technology. The circumstances of solder defect function are different for surface-mount technology

because the entire package passes through the wave. Solder joint skips are more common because the package body can shadow nearby leads or terminations from contact with the wave, causing an inadequate supply of solder at the joint. Therefore, the joint areas of the package should be fully exposed to the wave, as suggested by Fig. 2. Packages with leads or terminations on all four sides experience an increased propensity for defects on the leads in the nonpreferred orientation. Also, the greater density of devices on surface-mount circuit boards increases the frequency of solder skips due to the closer proximity of the packages. Finally, the finer lead pitch used on surface-mount leaded and leadless (chip carrier) packages limits or prohibits their assembly on boards by wave soldering. Process Parameters The second factor critical to successful implantation of wave soldering of printed circuit boards is the process parameters. Process parameters can be divided into three groups: the fluxing operation, the solder wave properties, and the process schedule. The fluxing operation not only includes the particular parameters (wave or foam height, spraying pressure, etc.) needed

to ensure adequate flux supply to the board for solder wetting, but also the condition of the flux. In-line fluxing techniques (wave or foam) frequently use the same quantity of flux to process a large number of boards. Deterioration of the flux due to board contaminants, as well as loss of the flux vehicle, must be monitored to minimize defects in the final product. Specific gravity is used to indicate flux condition. Low-solids fluxes are gaining popularity as a "no-clean" alternative to rosin-base fluxes, including their use in wave soldering. The low-solids contents (typically less than 5%) make these fluxes difficult to foam, so spray or wave applications are preferred (Ref 3). Tighter control of the flux chemistry, reduced solder pot temperatures (by 10 to 15 °C, or 18 to 27 °F), shorter solder dwell times (by 1 to 2 s), and the use of inert or reducing atmospheres (Ref 4) can reduce the number of defects. Solder Wave Properties. The second process factor is the solder wave properties. The simple wave geometry (also

called the T-wave) is shown in Fig. 1. Three regions of the wave are identified. The entry region (on the left with the board moving to the right) should be sufficiently turbulent to supply solder to all of the exposed metal surfaces and should permit the escape of flux volatiles that may cause voids in the joint. In the case of surface-mount assemblies, fluid motion must be sufficiently active to force solder close to the packages for adequate wetting of terminations and leads, thereby avoiding solder skips. The interior region of the wave acts as the heat source and solder supply to promote wetting of the substrate features. The exit or peelback region is the point at which the board leaves the wave and is critical in preventing solder bridge and icicle defects on the circuit board. For example, defects are reduced by having a smooth wave that separates very slowly from the board. This effect is achieved by matching the solder flow velocity with the speed of the circuit board (Ref 5) as well as by making adjustments to the wave configuration. Hot air knives positioned after the wave have been used to remove solder bridges and icicles prior to solidification of the solder. Figure 3 shows several equipment modifications used to produce different wave geometries. The extended T-wave and lambda-wave modifications to the standard T-wave cause a smooth surface at the exit region to limit bridges, icicles, and the formation of excessive dross. The dual-wave configuration provides separate turbulent and smooth (laminar) waves to effect the desired solder joint qualities. Note that these modifications generally require that the circuit board laminate and devices withstand a longer period of contact with the molten solder.

FIG. 3 EQUIPMENT MODIFICATIONS USED TO PRODUCE DIFFERENT WAVE GEOMETRIES. THE NOMENCLATURE DESCRIBES THE PROFILE GEOMETRY OF THE WAVE.

Dross formation tends to be high in wave soldering due to the large surface areas of solder exposed to the atmosphere (which is further aggravated by the turbulent flow of the solder in the pot). Dross particles can become embedded in the solder joints as well as disturb the peel back of the solder upon exiting the board. The use of inert atmospheres, reduced solder temperatures, and oils to cover the solder surface limit the formation of dross. Solder temperatures are typically higher than those used in reflow processes. The higher temperatures are required to provide sufficient heat transfer to the substrate for adequate wetting as well as to reduce the solder surface tension to limit defects at the peel-back region. Solder pot temperatures are typically 50 to 80 °C (90 to 145 °F) above the liquidus temperature of the alloy; higher values are required for thicker substrates (e.g., multilayer or controlled-expansion substrates). The potential for damage to the laminate, including warpage and distortion as well as delamination of the conductor, increases with higher solder temperatures. Continuous processing with wave soldering equipment can cause a build-up of contaminants in the solder pot (e.g., nickel, copper, and gold from circuit board and lead finishes). See the article "Soldering in Electronic Applications" in this Volume for guidelines on acceptability limits for solder pot contamination. In addition to the traditional tin-lead alloys, wave soldering has been adapted to other solders, including the eutectic tinbismuth solder and the tin-lead-bismuth alloys (Ref 6). Process Schedule. The third factor in the wave soldering technique is the process schedule, which includes the preheat

zone temperature profile, the solder wave temperature, the conveyor tilt angle (α in Fig. 1, which affects the exit geometry of the board), and the conveyor speed. The conveyor speed determines the dwell time of the circuit board in the preheat zone and solder wave. Although specific settings depend on the type of equipment and the product being assembled, some general guidelines can be applied. First, the preheat stage should bring the board temperature to the range of 100 to 150 °C (210 to 300 °F). These temperatures provide suitable activation of the flux and lessen thermal shock to the circuit board (and devices) upon contact with the solder wave. The dwell time in the molten solder wave should be limited to 3 to 5 s. Dwell times at the molten solder temperature can reach 5 to 10 s for dual-wave systems. Time-temperature exposure limits for soldering processes are typically specified for devices